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Bureau of Mines Information Circular/1988 



New Steelmaking Technology 
From the Bureau of Mines 

Proceedings of an Open Industry Briefing Held 
in Association With the Electric Furnace 
Conference, December 8, 1987, Chicago, IL 

Compiled by Staff, Bureau of IVIines 



i!(0^^^ 




UNITED STATES DEPARTMENT OF THE INTERIOR 




maJM^JjA^, 8//A^f 




~~ — ; 
Information Circular 9195 

w 



New Steelmaking Technology 
From the Bureau of Mines 

Proceedings of an Open Industry Briefing Held 
in Association With the Electric Furnace 
Conference, December 8, 1987, Chicago, IL 

Compiled by Staff, Bureau of Mines 



UNITED STATES DEPARTMENT OF THE INTERIOR 
Donald Paul Model, Secretary 

BUREAU OF MINES 
T S Ary, Director 




-r 






ui)' 



c\\ 



^5 



Library of Congress Cataloging-in-Publication Data 



New steelmaking technology from the Bureau of Mines. 

(Bureau of Mines information circular ; 9195) 

Bibliographies. 

Supt. of Docs, no.: I 28.27: 9195. 

1. Steel— Metallurgy— Congresses. I. Electric Furnace Conference (1987 : Chicago, IL). II. United 
States, Bureau of Mines. III. Series: Information circular (United States. Bureau of Mines) ; 9195. 



TN295.U4 



[TN730] 



622 s 



[669 '.1424] 



88-600072 



PREFACE 

On December 8, 1987, the Bureau of Mines held an open industry briefing in association with 
the Electric Furnace Conference sponsored by AIME's Iron and Steel Society. The papers presented 
at that briefing are contained in this Information Circular, which serves as a proceedings of the 
meeting. The papers highlight the Bureau's most recent research aimed at improving steelmaking 
technology. Areas addressed by this research include arc stability in electric steelmaking furnaces, 
stainless steel pickling processes, substitutes in steelmaking, steelmaking refractories, and recy- 
cling of steelmaking dusts and wastes. 

The open industry briefing used as a forum for the transfer of this research is one of the many 
mechanisms used by the Bureau of Mines in its efforts to move research developments, technol- 
ogy, and information resulting from its programs into industrial practice and use. To learn more 
about the Bureau's technology transfer program and how it can be useful to you, please write or 
telephone: 

Bureau of Mines 
Office of Technology Transfer 
2401 E Street, NW. 
Washington, DC 20241 
Telephone: 202-634-1224 



CONTENTS 



Preface i 

Abstract 1 

Introduction 1 

Improved Arc Stability in Electric Arc Furnace Steelmaking by Thomas L. Ochs and Alan D. Hartman 2 

Preheating of Ferrous Scrap by R. H. Nafziger and G. W. Elger 12 

Fluorspar Substitutes in Steelmaking by R. H. Nafziger and G. W. Elger 23 

Research on Basic Steelmaking Refractories by T. A. Clancy and J. P. Bennett 28 

Basic Research on Corrosion of Iron-Based Materials by David R. Flinn 33 

Fundamentals of Stainless Steel Acid Pickling Processes by Bernard S.Covino, Jr 39 

Decreased Acid Consumption in Stainless Steel Pickling Through Acid Recovery by G. L. Horter and J. B. Stephenson. ... 45 

Recycling of Stainless Steelmaking Dusts and Other Wastes by L. A. Neumeier and M. J. Adam 50 

Using Wastes as a Source of Zinc for Electrogalvanizing by V. R. Miller 58 

Economic Evaluation of a Technique To Pelletize Flue Dust and Other Waste From the Manufacture of Stainless Steel by 

Joan H . Schwier 67 





UNIT OF MEASURE ABBREVIATIONS USED IN THIS REPORT 


A 


ampere 


lb 


pound 


A/dm2 


ampere per square decimeter 


lb/ft3 


pound per cubic foot 


A/min 


angstrom per minute 


Ib/min 


pound per minute 


atm 


atmosphere, standard 


Ib/st 


pound per short ton 


at- pet 


atomic percent 


Mgal 


thousand gallons 


Btu/(lbmol) 


British thermal unit per pound per 


mg/(mincm^) 


milligrams per minute per square 




mole 




centimeter 


Btu/st 


British thermal unit per short ton 


min 


minute 


Btu/yr 


Brithsh thermal unit per year 


mL 


milliliter 


cm 


centimeter 


mm 


millimeter 


cm^ 


square centimeter 


MMBm 


million British thermal units 


°C 


degree Celsius 


MQ-cm 


megohm centimeter 


°F 


degree Fahrenheit 


mol/L 


mole per liter 


dB 


decibel 


ms 


millisecond 


dm2 


square decimeter 


m/s 


meter per second 


ft 


foot 


mt 


metric ton 


ft/s 


foot per second 


mV 


millivolt 


g 


gram 


nAlcrn^ 


microampere per square centimeter 


gal 


gallon 


/iP/cm^ 


microfarad per square centimeter 


g/cm^ 


gram per cubic centimeter 


nm 


micrometer 


g/L 


gram per liter 


lis 


microsecond 


g/m^ 


gram per square meter 


nm 


nanometer 


g-mol/dm^-h 


gram mole per square decimeter per 


Qcm^ 


ohm centimeter squared 




hour 


P 


poise 


gr/dscf 


grain per dry standard cubic foot 


pet 


percent 


h 


hour 


ppm 


part per million 


Hz 


hertz 


ppt 


part per thousand 


in 


inch 


psi 


pound per square inch 


K 


kelvin 


s 


second 


kcal/mol 


kilocalorie per mole 


scfm 


standard cubic foot per minute 


keV 


thousand electron volts 


st 


short ton 


kHz 


kilohertz 


st/d 


short ton per day 


kW 


kilowatt 


st/yr 


short ton per year 


kWh 


kilowatt hour 


V 


volt 


kW-h/lb 


kilowatt hour per pound 


V/h 


volt per hour 


kW-h/st 


kilowatt hour per short ton 


vol pet 


volume percent 


kW-h/yr 


kilowatt hour per year 


wt pet 


weight percent 


kVA 


kilovolt ampere 


yr 


year 


L 


liter 







NEW STEELMAKING TECHNOLOGY FROM THE BUREAU OF MINES 

Proceedings of an Open Industry Briefing Held in Association 
With the Electric Furnace Conference, December 8,1987, Chicago, IL 



Compiled by Staff, Bureau of Mines 



ABSTRACT 

This report is a proceedings of a briefing recently sponsored by the Bureau of Mines at which 
Bureau personnel presented findings from their research efforts to improve steelmaking techology 
currently used in the United States. The papers contained in this report address many areas of con- 
cern to the iron and steelmaking industry. Among these are improving arc stability in electric arc 
furnaces, preheating ferrous scrap to reduce energy consumption, fluorspar substitutes in steel- 
making, basic steelmaking refractories, corrosion of iron-based materials, improvements in stain- 
less steel acid pickling processes, recycling of stainless steel dusts and other wastes, and use of 
wastes as a source of zinc for electro/galvanizing. 



INTRODUCTION 



For over 50 years the Bureau of Mines has worked to improve 
technology used by a major constituent of the U.S. minerals 
industry — the iron and steelmakers. The most recent promising 
results of this research are presented in this report. Some of the 
research described has been completed; other research studies high- 
lighted are in progress. However, this report focuses on many sig- 
nificant findings with a probable high positive impact on the indus- 
try. For instance, the Bureau is studying the fundamental behavior 
of arcs in electric furnaces in order to improve the efficiency of 
this steel manufacturing technology. Past research has shown that 
electrical disturbances caused by the unpredictable arc are respon- 
sible for power surges and fluctuations and for noise levels in excess 
of 120 dB. Understanding fundamental arc behavior may enable 
researchers to control the arc and therefore optimize electric fur- 
nace operation and efficiency. Already these studies have yielded 
positive results in the identification of potential areas for change 
in furnace designs and operating procedures to greatly increase effi- 
ciency. 

In addition to this research, the Bureau has also evaluated the 
feasibility of preheating ferrous scrap charges in electric furnace 
operations to decrease energy consumption by using furnace off- 
gases. During a laboratory test, furnace offgases from a 1-st-capacity 
electric arc furnace were used to preheat continuously charged 
automotive scrap and metal stampings up to 1,110° F. Results 
showed that about 7 pet less electrical energy was used in charging 
preheated scrap than was used when cold scrap was charged. 

Other research conducted by the Bureau may help reduce costs 
associated with basic oxygen and electric furnace operations by 
providing a less expensive substitute for fluorspar fluidizers. Sub- 
stitutes evaluated as alternatives in basic oxygen flimace operations 



include colemanite, fused boric acid, synthetic fluorspar, and used 
aluminum smelter potlining. Alternatives for electric arc furnace 
operations were synthetic fluorspar, boric acid, hydroboracite, used 
aluminum poflining and anode tailing wastes, and Soreflux B 
(ilmenite). The substitute fluidizers did not adversely affect the steel 
produced in test operations. 

Along with optimizing the efficiency of fiimace operations, the 
Bureau is searching for ways to reduce the loss of strategic and 
critical metals during various phases of steelmaking and to reduce 
waste generation. This can be accomplished by recycling pickling 
solutions, wastes, and dusts. The Bureau has experimented with 
ion-selective membrane technology in developing a means to recycle 
acid solutions used during the pickling of stainless steels. Disposal 
of these acid solutions is costly to the steel manufacturing industry 
and results in the loss of valuable chromium and nickel. Prelimi- 
nary research has revealed that an electrodialysis cell using ion- 
selective membranes does have the potential for separating dissolved 
metals from spent pickling acid solutions while regenerating the 
acids for return to the pickling process. Through other research 
studies, the Bureau has developed a process permitting in-plant 
recovery of about 90 pet of the chromium, molybdenum, nickel, 
and iron from stainless steelmaking dusts and wastes. By using yet 
another Bureau-developed process, zinc extracted from electric arc 
furnace dust can be used for electrogalvanizing. 

Detailed accounts of the laboratory tests and results for each 
of these research studies are presented in this report. The report 
also provides a description of other studies conducted including an 
economic analysis of the technique to pelletize flue dusts and other 
waste resulting from steel manufacture in order to recover contained 
metals. 



IMPROVED ARC STABILITY IN ELECTRIC ARC FURNACE STEELMAKING 



By Thomas L. Ochs^ and Alan D. Hartman^ 



ABSTRACT 

In order to improve the performance of electric arc furnaces used to manufacture steel, the 
Bureau of Mines is studying the fundamental behavior of the electric arc in the electric arc furnace. 
Improvements in control and processes will allow more efficient and quieter operation of the elec- 
tric furnaces. Presently, electrical disturbances caused by the unpredictable arc are responsible 
for flicker and surges on the power grid and sound levels in excess of 120 dB in the vicinity of 
the furnace. It is these disturbances that the Bureau is investigating. 

In the Bureau experiments, electrical signals are sampled at 50,000 Hz and photographs of 
the arc are taken at up to 40,000 images per second using high-speed cinematography. These images 
are then correlated with the electrical signals to study the physical events in the arc plasma. Arcs 
studied to date indicate that the electrical waveforms have unique signatures preceding some volt- 
age excursions. Initial Bureau investigations indicate that mathematical techniques of analysis in 
the field of nonlinear dynamics have characteristics that enable these methods to speed up process- 
ing of the electrical signals from the arc for use as control parameters. 

The behavior of the arc in the experimental environment has led to the conclusion that there 
are new fiirnace design changes and operating procedures possible. The potential areas for change 
include the furnace shell geometry, continuous feeding methods, electromagnetic pumping of mol- 
ten metal, electrical control, furnace atmosphere, waste heat recovery, and electrode design. These 
changes could greatly increase efficiency, which is typically in the 60-pct range, improve furnace 
operation, and reduce noise. Interactions of these changes are complicated and must be considered 
together. 



INTRODUCTION 



The share of steel produced by electric arc furnaces has 
increased over the past 20 yr because of the flexibility of the minimill 
concept and the consequent reduction in costs to produce steel in 
the environment of rapidly fluctuating demand. There have been 
many changes in the arc furnace over the past 30 yr of use, includ- 
ing ladle metallurgy, ultrahigh power operation, water jacket cool- 
ing, oxyfuel burners, and scrap preheating. Control of the furnaces, 
however, has remained a initiative process based on the prior 
experience of the furnace manufacturer and the operator. This is 
because the high-current arcs used in electric arc furnace steelmaking 
are violent high-temperature conducting plasmas and have proven 
very difficult to understand over the past 100 yr of study. Because 
of this limited knowledge, the arcs are difficult to control (/, pp. 
15-16, 2). 3 Currents of 100,000 A are common in commercial arc 
furnaces. These currents can produce temperatures in the core of 
the arc of 12,000 to 15,000 K, which is more than twice as hot 
as the surface of the Sun. 



' Mechanical engineer. 
' Chemical engineer. 

Albany Research Center, Bureau of Mines, Albany, OR, 
' Italic numbers in parentheses refer to items in the list of references at the end 
of this paper. 



Although these arcs have been used in smelting and melting 
metals since the turn of the century, the fundamental processes taking 
place inside the arc and at the points where the arcs attach to 
the electrodes and the charge are poorly understood. Disrup- 
tions of these high-current arcs during operation can produce fluc- 
tuations on the power grid (3), ablation of the furnace refractory, 
and poor heat transfer to the melt. At the present time, control 
of the furnace is based on the experience of the operators and the 
control manufacturers, not on any fundamental understanding of 
the processes taking place in the arc. Presently, control systems 
react to reverse a past event (2) as opposed to acting to prevent 
a future event. 

Prevention of future events is possible only if the future situa- 
tion is predictable, based upon the past events. However, arc 
behavior presently is unpredictable. New methods of examining 
the signals available from the voltage and current waveforms in 
the transformer secondary circuitry, and therefore in the arc itself, 
may be able to supply unique signatures useful over the span of 
a wavelength for indicating disruptive events. If the arc and its inter- 
actions with the furnace interior were better understood, then there 
could be improved control methods or modified equipment designs 
that would result in gains in efficiency and a reduction of disrup- 
tive electrical and acoustical noise. 



The electric arcs studied in the Bureau investigations have 
shown deterministic behavior that is sensitive to the conditions of 
operation. Extreme sensitivity to operating conditions leads to a 
lack of predictability of the arc behavior since the furnace operat- 
ing conditions cannot be measured exactly. This unpredictable arc 
behavior has been characterized as stochastic, or random, in prior 
studies, but instead is indicative of chaotic behavior resulting from 
nonlinear interactions. Viewing the arc as a chaotic, deterministic 
system of discrete events, it is possible to look at electrical wave- 
forms and expect short-term precursors (half-cycle) to the seem- 
ingly random events. Short-term precursors indicate the possibility 



of anticipation and control of arc disruption. It is disruption of the 
arc while it is carrying a high current that causes both electrical 
and acoustical noise. 

From these investigations, it is becoming clear that traditional 
transform methods of analysis and statistical analysis of the elec- 
trical waveforms are of limited use. These methods are normally 
used on smooth waveforms or waveforms with a small number of 
periodic discrete events. The waveforms that were obtained in this 
research are composed of many nonperiodic discrete events on dis- 
torted square and sine waves. These discrete events must be treated 
using discrete digital methods. 



STOCHASTIC VERSUS CHAOTIC BEHAVIOR 



The electric arc has been described as a random, or stochas- 
tic, system, with events taking place at unpredictable intervals. The 
present investigation shows that the arc is not a stochastic system, 
but rather a nonlinear system that is very sensitive to initial condi- 
tions. This sensitivity gives the appearance of random behavior since 
the internal conditions cannot be measured closely enough to pre- 
dict the next operational state of the arc (4-5). Behavior of these 
nonlinear chaotic systems is not possible to predict for any length 
of time from a mathematical solution based on measured system 
conditions. Instead, the system can be treated using the methods 
of nonlinear dynamics. Using these techniques, there is the possi- 
bility of short-term (one ac cycle) prediction of future events based 
upon the inferred conditions as deduced from the signature analy- 
sis of the real-time waveforms. 



The basic premise of this method of predictive control is that 
the system is mathematically well behaved (continuous and single 
valued), and over a short term (one cycle) the arc behavior can be 
predicted. Over the long term (longer than a cycle), the minor var- 
iations in operating conditions that cannot be measured will pro- 
duce unpredictable behavior even though the system is deterministic. 
This means that the control system must take real-time data of the 
arc waveforms, compare it against a library of waveform signa- 
tures, and make decisions in a short time frame, typically about 
a quarter-cycle (4 ms). Recent increases in computational capabil- 
ities and decreases in cost have made data analysis and system con- 
trol of the type described feasible. 



EQUIPMENT 



An experimental 200-lb-capacity single-phase electric arc fur- 
nace was used for conducting the experiments. The furnace used 
two 3-in-diam graphite electrodes. The power was supplied by two 
single-phase ac welders connected in parallel. Each welder was rated 
at 1,500-A current and had a rated full-load voltage of 40 V. The 
primary rating was 440 V and 170 A single-phase. 

The furnace was modified in order to simplify the data analy- 
sis. Three types of arc targets were used in the furnace. The arc 
target block materials were graphite, steel, and copper. These blocks 
were used to study different system configurations and obtain calori- 
metric data. The calorimetric data are for assessment of heat transfer 
rates and efficiency. Initially, electrical signals in the single-phase 
furnace were taken across both arcs, one from each of the elec- 
trodes. In this configuration, the two arcs consisted of one with 
the electrode acting as the cathode and one with the electrode act- 
ing as the anode. This averaged the events attributable to each arc 
and made data analysis difficult. Therefore, a simplifying modifi- 
cation was made to the experimental system. This consisted of 
threading one of the electrodes into the conductive target block (fig. 
1). Threading the electrode into the block eliminated one of the 
arcs and its associated signals, while it maintained the current path 
that normally would be taken by the current flowing through the 



arc. This made the magnetic field in the ftimace similar to that pres- 
ent in the two-arc system. 

The second modification to the experimental furnace involved 
replacing the furnace shell with an airtight enclosure. By using this 
enclosure, the atmosphere within the ftimace could be controlled 
(fig. 2). This allowed experimentation to be conducted with gases 
other than air, as well as gas injection through the electrode. 

A third modification was the addition of two viewports at 90 ° 
to each other (fig. 3). Thus two perpendicular images of the arc 
could be captured simultaneously by a high-speed motion picture 
camera. The two images of the arc were directed into the lens of 
a high-speed camera by the use of mirrors. The camera used 450-ft 
rolls of 16-mm film and operated at up to 1 1,000 frames per sec- 
ond. The actual image capture can take place at up to 44,000 image 
pairs per second, which was accomplished by using the camera's 
internal prism to divide a frame into quarters (fig. 4). After 200 
ft of film had been exposed, allowing the camera to reach maxi- 
mum speed, a waveform analyzer was triggered that recorded the 
simultaneous voltage and current electrical signals corresponding 
with the film images. Synchronization between the waveforms and 
the film is achieved by the use of timing pulses on the film. The 
waveforms were digitized at 50 kHz per channel. 



r'^ 







»', 




Figure 1.— Electrode threaded into target block to simplify target path. 




Figure 2.— Experimental furnace shell showing gas Inlets. 




Figure 3.— Orthogonal views by use of mirror system. 



Rotating prism assembly 
Aperture mask 

-Objective lens 



Film frame at film 
gate aperture - 




Viewf inder eyepiece 
Focusing prism 



First field lens and 
prism assembly 



Figure 4.— View of single frame showing eight images and their relationship to original image. 



EXPERIMENTAL PROCEDURE AND RESULTS 



Areas investigated by the Bureau include (1) electrode tip 
design, (2) inert gas injection, and (3) arc analysis. Electrode tip 
design was investigated since the structure of the electrode provid- 
ing the arc has been related to the unsteadiness of arcs (6). 



ELECTRODE TIP DESIGN 

Electrodes providing a concave tip were found to reduce arc 
flare and maintain the arc under the electrode tip (7). In reference 
7, the author described attempts to reproduce this effect by testing 
hollow versus solid electrodes. Although the hollow electrodes 
increased the heating efficiency of the test furnace by approximately 
10 pet over solid electrodes, the hollow electrode consumption was 
2 to 10 pet above that of the solid electrodes. In order to improve 
energy efficiency while reducing electrode consumption, the Bureau 
designed alternative electrode tips. A basic electrode tip design used 
in the Bureau's research was a 1-in-diam tip machined onto the end 
of the 3-in-diam electrode (fig. 5). The theory behind this tip design 
is that the arcing between the workpiece in the furnace and the elec- 
trode will remain concentrated onto the 1-in-diam tip. The tip will 
become heated while the remaining 3-in-diam section of the elec- 
trode will remain relatively cool. This phenomenon was demon- 
strated in the high-speed films of the arc since the tip glowed white 
hot while the remaining, larger diameter electrode section surface 
remained black. 

By making the smaller diameter electrode tip section an expend- 
able section, the larger outer section of the electrode could become 
a semipermanent structure, thus reducing the amount of graphite 
needed for arcing (fig. 6). 



INERT GAS INJECTION 

Inert gas injection was investigated to decrease electrode con- 
sumption by replacing the oxidizing atmosphere with an inert 
atmosphere. The inert gas was introduced into the furnace at three 
locations, through each viewport and also through the arcing elec- 
trode. Holes 1/32 in. in diameter on a 1-in-diam circle encompass- 
ing the electrode tip were used to introduce the gas (fig. 7). The 
gas from the 15 exit holes shrouded the arc and maintained it in 
a vertical direction, thus allowing more of the heat to be directed 
to the melt. The gas also helped to confine the arc to the tip section 
while it cooled the outer portion of the electrode, again decreasing 
electrode consumption. 



ARC ANALYSIS 

Electrical waveforms of voltage and current were monitored 
across the arc with a waveform analyzer. The waveforms differed 
markedly depending upon the target composition. The differences 
between arc waveforms when using graphite, steel, and copper tar- 
gets, are easily visible (fig. 8). These differences are expected since 
the thermal and electronic properties, such as the melting point and 
the amount of energy needed to free an electron from the surface 
of a material, vary dramatically from material to material. Figure 
8 shows the voltage as a solid line and the current as a dashed line. 
The spikes on the voltage waveforms corresponded to the move- 
ment of the arc as seen in the high-speed films. In figure 8 (top), 
the positive and negative half-cycles have roughly the same abso- 
lute amplitude because the arc occurred between a graphite elec- 
trode and a graphite target. However, in the center and bottom 





R 1-1/2 



R 1/2 



TOP VIEW 



SIDE VIEW 



Figure 5.— Button electrode tip. 




Feedable 
portion of 
electrode 



Stationary 
electrode 

shell 



Figure 6.— Center feed electrode tip. 




R 1/16' 




TOP VIEW 



SIDE VIEW 
Figure 7.— Button electrode tip with gas injection. 



panels of figure 8, the positive and negative half-cycle amplitudes 
show asymmetry owing to arcing between a graphite electrode and 
steel or copper, respectively. This asymmetry is due to the ability 
of graphite to emit electrons thermionically, whereas metals melt 
before thermionic emission occurs. In each of the waveforms, the 
electrode is acting as the anode in the positive half-cycles. 

Waveforms and the corresponding high-speed films were ana- 
lyzed together to identify the arc characteristics that were associated 
with the millisecond events on the waveforms. An example is shown 
in figure 9. In this case, the furnace atmosphere was 100 pet Ar 
and the arc was between a graphite electrode and a copper block. 
In the figure, voltage is the solid waveform and current is the dashed 
waveform. Positive half-cycles 1 and 3 are at a relatively low volt- 




60 
< 


1 1 


N 
O 




^ 40 


— 




— 


1— " 


Voltoqe-i 








Z 


Current^ 








a: 
a: 20 

3 


n 


\ 


^ 


^ 


r-\ 


r^ 


- 


O 

T3 

C 

O 


^--^^ 


_ 


K 


V "/ 


,/.:-> 


\ — "/ 


^--> 


K"---/ 


r::> 


\ 


> 

LJ 
O-20 




V V 




1— 




o 




> 


1 1 



80 




TIME.s 

Figure 8.— Waveform of voltage and current for arc from graph- 
ite electrode to graphite (top), steel (center), and copper (bot- 
tom) target in 90-pct-He, 10-pct-Ar atmosphere. 



age as compared to positive half-cycles 2 and 4. Half-cycles 1 and 
3 could be matched to short arcs between the graphite tip and cop- 
per block as the arc target, whereas half-cycles 2 and 4 correspond 
to long arcs on the film that were between the 3-in-diam section 
of the electrode and the copper block arc target. 

Many shorter term discrete events of much less than a half- 
cycle duration have been identified in both voltage and current wave- 
forms. In figure 10, four major events on a single-cycle voltage 
waveform have been correlated with the arc movement in the high- 
speed films. These events took place in a 90-pct-He, 10-pct-Ar 
atmosphere. Event A, which is a positive voltage spike, cor- 
responded to the arc changing positions from arcing a short dis- 
tance from the electrode tip, to arcing a longer distance, between 
the 3-in-diam section of the graphite electrode and the copper block 
as the arc target. The maximum voltage for event A was 53 V and 
occurred as the arc developed into the long arc column as depicted 
in sequence A of figure 10. Event B, on the positive side, showed 
that the rapid fluctuation of the voltage related to a long arc decreas- 
ing to a short arc on the tip section of the electrode and then back 
again to a long arc for each fluctuation. Event C, termed a shark's 
tooth, on the negative half-cycle could be related to the arc's move- 
ment across the tip section in a right-to-left motion, and finally event 
D was the opposite motion of the arc moving from the left to the 
right side on the electrode tip section. 

A similar experiment with an atmosphere of 90 pet He- 10 pet 
Ar produced the arc motions as shown in figure 1 1 . Five areas are 
related to the arc's movement in the high-speed films. Area 1 related 
to a short arc rotating directly beneath the electrode's tip. Area 2 
showed the gradual movement of a long arc, between the 3-in-diam 
graphite electrode and copper block, from the tip to the outer edge 
of the electrode. Area 3 is a short, very stationary arc on the nega- 
tive half-cycle of the voltage waveform. Area 4 was correlated with 
the quick movement of the short arc from the left side to the right 
side on the tip. Area 5 was a rotating short arc directly beneath 
the electrode tip. 




B 



Long arc 



± 



T 



I 



Copper block 

Figure 9.— Waveform. A, Long arc; B, short arc. 




6,680 10.020 

TIME, ps 



Area 


1 1 

Area 


1 




' 


1 

r 


A_^ 2 






- 


- 




1 Area 


Area 


Area 




1 3 
1 1 1 


4 


5 

1 



3,320 



6,640 9,960 

TIME, ;js 



16,600 




Rotary 
motion 




Plasma W 



Copper block' 



%r 




Arc motion 




^ 



1 r 




11 Direction |__J 
X^ mo veme nt ^{) 



Figure 10.— One cycle with discrete events (A-D) delineated 
and diagramed. 







ovement 



Figure 1 1 .—One cycle with five discrete areas delineated and 
diagramed. 1, rotating, short arc; 2, arc moving right to left, arc- 
ing between 3-in-diam section and copper block; 3, very station- 
ary arc; 4, arc movement; 5, rotating arc under tip. 



10 



Image analysis is being used to help define the arc's core within 
the plasma shown on the high-speed film ft-ames. The arc structure 
is difficult to resolve because of the high luminosity that tends to 
saturate photographic emulsions, producing a seemingly white arc. 
Image analysis is helping to resolve the inner structure by digitally 
filtering the images to reconstruct the actual radiant intensity gra- 



dients. From these intensity gradients, the electron paths can be 
used to relate arc parameters to the resistance of the arc. For 
instance, image analysis is used to map areas of highest light inten- 
sity so that an arc path of conduction can be defined. By using this 
technique it has been possible to measure an arc length (fig. 12). 



SYSTEM INTERACTION 



The occurrence of a break in the arc at high current or the short- 
ing of scrap to the electrode is responsible for electrode breakage, 
increased melt times, and voltage spikes that can damage electri- 
cal equipment, cause flicker on the electric power grid, and cause 
excessive noise. The identified events that cause these disruptions 
of the furnace arc are related to motion of the arc, scrap, electrodes, 
and/or magnetic fields, as seen in the Bureau experiments. When 
methods of stabilizing the arc in the furnace environment are con- 
sidered, it is necessary to consider all of the interactions between 
the furnace variables, such as slag composition, atmosphere com- 
position, type of scrap, temperature of the melt, composition of 
the electrodes, electrode geometry, and control methods. Early in 



this investigation, it became clear that there is a synergistic rela- 
tionship between the variables in the furnace interior that is a direct 
result of the nonlinear relationships of the coupled magneto- 
hydrodynamic equations governing the electric arc behavior. 
Because of this interactive behavior, it is not possible to change 
one variable without affecting the other furnace variables. Presentiy, 
most of the important parameters in the electric furnace are allowed 
to float at whatever value they may take. For example, there is no 
mechanism to control furnace atmosphere composition, instantane- 
ous voltage, or geometry of the electrodes. These variables have 
been shown in experimentation to be very important for arc 
operation. 



ATMOSPHERE CONTROL 



One of the most influential variables in the arc experiments 
is the furnace atmosphere. Diatomic molecules such as oxygen and 
nitrogen must be disassociated before they can be ionized, and there- 
fore, they are more difficult to ionize than inert gases. If inert gases 
are used, then the electrode consumption is dramatically decreased 
since there is no oxygen or nitrogen to react with the graphite. These 
reactive diatomic gases also will react with the steel if they are pres- 
ent, and if they are absent, then the steel composition can be more 
closely controlled. The gases found to be most promising in this 
study are mixtures of helium and argon. These mixtures have good 
heat transfer properties and are easy to ionize. 

It has also been indicated in the studies of the arc fluid dynamics 
that shrouding the arc in a flowing gas will help to stabilize the 



arc as in a plasma torch. This comes about by a combination of 
a thermal pinch (contraction of the arc due to cooling of the outer 
arc surface and a subsequent reduction in electrical conductivity, 
forcing the electron flow into the center of the arc), causing wall 
stabilization and actual fluid dynamic forces preventing the arc from 
migrating through the gas shroud (fig. 13). This indicates that 
properly engineered gas injection through the electrode would help 
to stabilize the arc. This plasma jet effect also will increase con- 
vective heat transfer to the melt, thereby increasing thermal effi- 
ciency. The nonreactive inert gases also will permit the use of a 
wide range of slags that cannot be used in the traditional air 
atmosphere, and the gas injection through the electrodes possibly 
could be used to introduce reductants as needed. 




Figure 12.— Processed image (actual size) showing arc path 
between 3-ln-diam section of graphite electrode (upper right- 
hand attachment point) and graphite block as the arc target (near 
bottom of 1-in button on electrode). 




Figure 13.— Gas shroud around transferred arc showing forces 
tending to stabilize arc (actual size). 



11 



CONTROL SYSTEM 



A computerized control system could be operated by using a 
data base of historical waveform information and pattern match- 
ing information to determine the real-time furnace operating con- 
ditions. Using the data such as voltage and current waveforms, 
atmosphere composition, feed material, and slag type, the control 
system could adjust the furnace operating conditions for optimal 
performance. The adjustments would include variables such as 
atmosphere composition, feed rate, feed composition, voltage, gas 
injection rate, magnetic stirring, and electrode position. These 
adjustments could be made at rates that depend on the parameter 
being adjusted. For instance, if the parameter is voltage, then the 
adjustments must be made in the course of a half-cycle (8 ms). On 
the other hand, if the parameter is feed material, then the adjust- 
ment will be allowed to take place over a period of minutes. For 
indicators of a catastrophic disruption such as breaking of the arc 



during high-current conditions, the most simple control strategy 
would be to turn the current off at a zero current crossing, then 
position the electrodes to restart seconds later. More complex con- 
trol would involve electronic tap changing and electrode control 
to maintain the arc under adverse conditions. This would prevent 
disruption at high current. 

The type of system best suited to this process of decisionmak- 
ing based on data and experience is an expert system. The system 
could maintain a knowledge base of waveforms and past experience. 
Over a period of time, the system can be customized to the individual 
furnace it operates on by logging any uncataloged events and the 
corresponding results so that future decisions could be made based 
on this experience. These expert systems are beginning to be seen 
throughout industry . Adaptive expert systems are now being devel- 
oped and within the next few years will be corrmiercially available. 



SUMMARY 



A new perspective is available for investigation of electric arc 
behavior through the use of high-speed motion pictures and syn- 
chronized electrical waveforms. Analysis of the seemingly random 
occurrences in the arc on an event-by-event basis shows that the 
arc is deterministic and hence theoretically controllable. High-speed 
computers now make it possible to economically control factors 
such as transformer tap settings, furnace atmosphere, and electrode 



positions. These control factors coupled with new electrode geom- 
etry can improve efficiency and yield while stabilizing the elec- 
tric arc. 

New ways have been developed for investigating electric arc 
behavior through the use of high-speed motion pictures synchronized 
with electrical waveform data. 



REFERENCES 



1. Ochs, T. L., A. D. Hartman, and S. L. Witkowski. Waveform Anal- 
ysis of Electric Furnace Arcs as a Diagnostic Tool. BuMines RI 9029, 1986, 
19 pp. 

2. Paschkis, M. E., and J. Persson. Open-Arc Furnaces. Ch. in Indus- 
trial Electric Furnaces and Appliances. Interscience, 2d ed., 1960, pp. 
179-228. 

3. Schwabe, W. E. Arc Furnace Power Delivery Scoping Study. Elec- 
tric Power Res. Inst., Palo Alto, CA, EPRI RP-1201-24, 1982, 146 pp. 

4. Abraham, R. H., and C. D. Shaw. Chaotic Behavior. Part 2 of 
Dynamics, The Geometry of Behavior. Aerial Press Inc., Santa Cruz, CA, 
1983, pp. 99-105. 



5. Crutchfield, J. P., J. D. Farmer, N. H. Packard, and R. S. Shaw. 
Chaos. Sci. Am., v. 255, No. 6, 1986, pp. 46-57. 

6. Schwabe, W. E. Lighting Flicker Caused by Electric Arc Furnaces. 
Iron and Steel Eng., Aug. 1958, pp. 93-100. 

7. . Experimental Results With Hollow Electrodes in Electric Steel 

Furnaces. Iron and Steel Eng., June 1957, pp. 84-92. 



12 



PREHEATING OF FERROUS SCRAP 



By R. H. Nafzigeri and G. W. Elger^ 



ABSTRACT 



Energy conservation is an important consideration in all steelmaking operations. Energy con- 
sumption impacts on the productivity and costs of producing steel. Accordingly, the Bureau of 
Mines has evaluated the feasibility of preheating ferrous scrap charges in basic oxygen furnace 
(BOF) and electric furnace steelmaking operations to decrease energy consumption. Offgases gener- 
ated during oxygen blowing of a molten charge in a !4-st-capacity BOF were passed through a 
static bed of shredded auto scrap. Final bed temperatures ranged from 1,150° to 1,650° F. The 
thermal energy recovered can contribute up to 44 pet of the energy necessary to melt the scrap. 
Furnace offgases from a 1-st-capacity electric arc furnace were used to preheat continuously charged 
automotive scrap and metal stampings up to 1,1 10° F. Approximately 7 pet less electrical energy 
was used compared with that consumed in continuously charging cold scrap. Conventional back- 
charging techniques also were used for comparison purposes. 



INTRODUCTION 



In 1985, electric arc furnaces produced 34 pet or nearly 30 
million st of raw steel in the United States (1).^ Most of the 
remainder (59 pet) was produced in BOF's. Electric furnace steel- 
making operations use cold ferrous scrap nearly exclusively as 
charge materials, and a considerable amount of cold scrap is used 
by the BOF. Energy consumption is one of the primary concerns 
and a major cost in domestic steelmaking operations. For exam- 
ple, approximately 535 kW-h/st of steel produced is required in 
electric arc furnace steelmaking (2). This represents approximately 
1.8 X 10* Btu/st of steel. The fuel and electrical energy consump- 
tion in a BOF is approximately 0.9 x 10* Btu/st (3). This represents 
1.14 X 10'" Btu/yr consumed in domestic steelmaking. A 1-pct 
decrease would save 1.14 x 10'^ Btu/yr or 334 x 10* kWh/yr. 

In both types of steelmaking operations, the Bureau is striving 
to increase the recycling of scrap. The BOF experiments were con- 
ducted to assess the feasibility of increasing the proportion of scrap 
used in the charge mixture. In addition, the use of hot offgases for 
preheating offers the potential of decreasing energy consumption 
and costs. The additional amount of scrap used could offset the 
limited availability of hot metal owing to blast furnace shutdowns, 
for example. The evaluation of preheated scrap in electric arc fur- 
nace steelmaking was aimed at promoting scrap use by making elec- 
tric arc furnace steelmaking more competitive through decreased 
electrical energy consumption and costs. Early Bureau research 
involved the use of waste heat in electric arc furnace offgases to 
preheat prereduced iron ore pellets during continuous charging of 
the furnace (4-5). 



' Research supervisor. 
2 Research chemist. 

Albany Research Center, Bureau of Mines, Albany, OR. 
^ Italic numbers in parentheses refer to items in the list of references at the end 
of this paper. 



Others have discussed the preheating of scrap in BOF opera- 
tions. Several methods have been described, including (1) in- vessel 
preheating, (2) separate vessel preheating with the thermal energy 
provided by oxygen-natural gas or oxygen-fuel oil burners (6-7), 
and (3) waste gas preheating with the heat derived from offgases 
generated during oxygen blowing of the BOF charge (8). In one 
application, natural gas-oxygen burners preheated scrap charges 
prior to the molten iron addition. Increased productivity and lower 
ingot costs were realized, but excessive scrap oxidation was noted 
(9). In another study, increased scrap could be charged to a BOF 
when it was preheated to 1,700° F. This decreased lime and flux 
consumption, decreased slag volume, and decreased metal blow- 
ing time. However, refractory consumption was increased (]0). 

Developments in preheating scrap charges for electric furnace 
steelmaking occurred as early as 30 yr ago. In nearly all cases, some 
means of external heating was used. Three techniques for preheat- 
ing have been used. The first utilizes a special vessel for preheat- 
ing. Typically, external fuel burners supplement the heat from the 
offgases used. After preheating, the scrap is transferred to a charging 
bucket prior to introduction into the furnace (11). Considerable scrap 
handling, large space requirements, and high costs for the vessel 
pit are cited as disadvantages. 

A second method involves placing the charge bucket into a pre- 
heating vessel. In this case, the scrap cannot be preheated to a high 
temperature, and there is more dust (12). Variations in this tech- 
nique involve the direction of hot furnace offgases to preheating 
stands or to a preheating chamber to heat the scrap in a charging 
bucket (13-14). 

In the third technique, a bucket lined with castable refracto- 
ries serves as a preheating vessel and a charging bucket. Potential 
bucket distortion, dust losses, and a weight that requires a high crane 
capacity can cause problems with this method. 



13 



Oil burners either mounted in the charging bucket or in the 
furnace, natural gas lances, the use of charging buckets with lou- 
vers placed over hot billets or ingots, or the use of preheat cham- 
bers are additional preheating techniques that have been used 
(15-22). 

All of these techniques involve backcharging methods for feed- 
ing the furnace. Relatively high capital costs have precluded the 
adoption of fuel-fired preheaters. Other disadvantages include (1) 
distortion or warpage of the charging bucket doors, (2) uneven heat 
distribution within the charge, and (3) scrap oxidation. 



Objectives of the Bureau research reported herein include 
(1) an evaluation of preheating BOF scrap charges using recovered 
hot offgases that are passed through a scrap bed in a separate cham- 
ber to eliminate oxygen-fuel preheating, and (2) a determination 
of the feasibility of continuously charging fragmented scrap into 
an electric arc furnace whereby the scrap is heated by countercur- 
rent hot furnace offgases to realize decreased electrical energy 
requirements compared with those necessary in conventional back- 
charging tests. 



BOF EXPERIMENTS 



EXPERIMENTAL EQUIPMENT, MATERIALS, 
AND PROCEDURES 

All of the tests were performed in a 'i-st capacity BOF, shown 
in figure 1 . Offgases generated during oxygen blowing of the fur- 
nace charge were drawn through an adjacent preheat vessel by a 
blower located downstream from the preheater in the dust collect- 
ing system. The preheat vessel contained the scrap charge. A 
schematic diagram of the system is depicted in figure 2. Further 
details have been published previously (23-24). 

The furnace charge consisted of shredded automobile scrap hav- 
ing pieces no larger than 3 by 3 in, with a bulk density of 92 lb/ft\ 




This material was fed into the preheat vessel prior to blowing the 
BOF metallic charge. After the blow, the heated scrap was used 
in the next test as a replacement for the normal cold scrap charge. 
Offgas flow was controlled by a manual slide damper located down- 
stream from the preheater. The BOF charges contained 100 to 180 
lb of preheated automobile scrap and 270 to 350 lb of molten pig 



Damper 



Duct 



Lance 




Scrap 
preheater 



Insulated duct 



Basic oxygen 
furnace 



Figure 1 . — Basic oxygen furnace and preheater system. 



Figure 2.— Schematic diagram of the BOF-preheater system. 



14 



iron. Typical charge compositions are summarizeci in table 1 
(23-24). 

After the preheated scrap-hot metal mixtures were charged, 
the oxygen lance was positioned, and blowing began after 25 lb 
of lime and 1 lb of fluorspar were added to the furnace. Oxygen 
blowing, at a fixed rate of 27 scfm, was terminated when visual 
observations of flame height and color indicated a carbon level of 
0. 10 pet. Blow times varied from 12 to 17 min and were depend- 
ent upon the quantity of scrap added. 



RESULTS 

An initial series of tests showed that cold scrap could consti- 
tute up to 28 pet of the charge. Cold scrap additions above this level 
resulted in a significantly lower oxygen efficiency and deteriorat- 
ing furnace operations (23-24). 

A second series of experiments were conducted in which pre- 
heated scrap charged to the BOF ranged from 22 to 40 pet of the 
total charge. Data from these tests are shown in table 2. Average 
preheated scrap temperatures ranged from 1,050° to 1,650° F, with 
the lower amount of preheated scrap resulting in the highest scrap 
temperature. Scrap temperatures were dependent upon the quan- 
tity of scrap in the preheater and upon the length of the oxygen 
blowing period. With 40 pet preheated scrap, the average scrap 
temperature increased as a result of a significantly lengthened blow 
time with lower oxygen efficiency. At 1,650° F, sufficient heat 
is stored in the scrap to yield approximately 44 pet of the energy 
required for melting. Results indicated that a 40-pct preheated scrap 
charge was the maximum that could be tolerated by the Bureau's 



Table 1 .—Typical chemical analysis of BOF metallic charge 
and product, weight percent 



Description 


C 


Cu 


Mn 


P 


Si 


Scrap .... 
Hot metal. 
SteeM .... 


<0.1 
4.0 
0.12- .15 


0.12 
.25 

.21 


0.062 
.80 

:5.1 


0.011 
.07 
^.01 


0.32 
1.0 

^.1 



Steel product from 33-pct scrap addition. 



Table 2.— Average temperature and heat recovery values of 
preheater scrap charges 



Preheated Av scrap Heat 


Total heat 


Heat 


scrap, temp, content, ' 


recovered, 


recovered. 


pet of °F Btu/(lb'mol) 


Btu 


pet required 


BOF charge 




for melting^ 


22 1,650 14,240 


25,500 


44 


27 1,265 9,770 


21,000 


30 


29 1,175 8,630 


20,100 


27 


33 1 ,045 7,330 


19,700 


23 


36 1 ,050 7,380 


21,100 


23 


40 1,150 8.380 


27,000 


26 


' From BuMines B 476, 1949, p. 85. 






Heat content, Btu/db-mol) 






2 Calculated = oo ,o., o. .»,k.' u 


X 100. 





BOF to maintain satisfactory oxygen efficiency and sufficient metal 
temperature for tapping. Therefore, preheating the scrap increases 
scrap utilization from 28 pet of the charge to 40 pet, an increase 
of 43 pet. On the basis of these tests, 36 pet of preheated scrap 
appeared optimum with respect to final steel temperatures ( = 3,000° 
F) and oxygen efficiency. Typical temperatures of preheater 
entrance and exit gases and scrap under these conditions are shown 
in figure 3 (23-24). 

Scrap preheating was enhanced by the oxidation of the CO, 
formed in the BOF, to CO2 before reaching the preheater. This was 
caused by secondary air infiltration around the BOF hood. Signifi- 
cant scrap oxidation occurred only when scrap bed temperatures 
exceeded 1,800° F. A tighter fitting hood over the BOF would 
decrease the air infiltration and perhaps allow higher offgas inlet 
temperatures in the preheater without excessive scrap oxidation 
(23-24). However, the heat content of the inlet gas would be lower 
and the scrap preheat temperature therefore would be lower. 



1- 
< 

cr 

UJ 

a. 

UJ 



2,550 



2,250 



L950 - 



.650 



1,350 



1,050 



750 



450 



I I I I ' I r 

KEY 
A Offgas enfrance temp 
• Average scrap temp 
■ Offgas exit t( 




4 8 12 16 

OXYGEN BLOWING TIME, min 



20 



32,130 Btu/(lb-mol) 
(Denominator is the heat content of steel at 2,786° 



F.) 



Figure 3.— Effect of oxygen blowing time on preheater offgas 
and scrap temperatures. Charge condition: BOF = 160 lb scrap 
+ 290 lb hot metal, preheater = 160 lb scrap. 



15 



ELECTRIC ARC FURNACE EXPERIMENTS 



EXPERIMENTAL EQUIPMENT, MATERIALS, 
AND PROCEDURES 

All tests were conducted in a conventional electric arc steel- 
making furnace of 1-st capacity, lined with a basic brick and cov- 
ered with a rammed-alumina roof, as shown in figure 4. The 
electrical energy was provided by a 1,200-kVA transformer 
through three 4-in-diam graphite electrodes. A stainless steel chute 
was attached to the furnace. At the opposite end of the chute was 
a rotating feeder-preheater that consisted of six 12-in-diam sections 
of stainless steel tubing connected at 90 ° to each other in a zigzag 
fashion. Figure 5 shows a schematic diagram of the furnace, scrap 
feeder-preheater, and furnace offgas ductwork, including gas sam- 
pling location and dust removal units. In use, the charge was fed 
through the chute and preheater, with the hot exhaust gases pass- 
ing through the preheater (3)* countercurrent to the direction of 
the charge, thereby heating the charge and cooling the gases. 

The offgases exited directly into a long vertical section of duct 
that assisted in the removal of those smaller sized dust particles 
escaping a stainless steel cyclone. The gases then were directed 
horizontally to a second cyclone (1) and an adjacent baghouse for 
final cleaning before being exhausted to the atmosphere. Gas and 
particulate samples were taken in this horizontal section (2). This 
furnace has been described previously in more detail (5, 25). The 
components in the sampling train for particulate stack sampling. 



" Bold numbers in parentheses refer to components identified in illustrations. 




shown in figure 6, included the probe with attached nozzle (1), a 
particulate filter (4), a cooling and/or gas collector with four 
impingers (5), flow-measurement devices, and a vacuum pump (13). 
Other components shown in figure 6 include a cyclone (2), flask 
(3), thermometers (6), check valve (7) connecting cord (8), vacuum 
gauge (9), coarse adjust valve (10), fine adjust valve (11), oiler 
(12), filter (14), dry-gas meter (15), orifice tube (16), incline 
manometer (17), solenoid valves (18), pitot (19) thermocouple (20) 
and temperature recorder (21). This equipment and the procedures 
used for gas sampling also have been described in more detail previ- 
ously (25). 

The shredded scrap used for these tests was purchased from 
a local scrap processor and consisted of three separate batches, each 
purchased at a different time and with a different composition. The 
metal stampings used in the charges were purchased from the same 
source. Only pieces of scrap with largest dimensions of less than 
4 in were used. The chemical analyses of these materials are shown 
in table 3 . The analyses were obtained by a direct-reading spectro- 
graph for cast samples melted in separate 800-lb wash heats with- 
out quartz and lime additions to provide a slag cover. Scrap 
meltdown without a slag cover was conducted to keep the alumi- 
num in the metal phase rather than transfer that constituent to the 
slag phase. 



KEY 
/ Exhaust gas cleaner 
2 Gas sampling 
J Preheater 
4 Electric-arc furnace 




Figure 4.— One-short ton electric arc furnace and feeder- 
preheater used for steelmaking tests. 



Figure 5.— Schematic diagram of electric arc furnace, feeder- 
preheater, and offgas ductwork (not to scale). 



16 



Table 3. — Chemical composition of shredded automotive 
seraph and metal stampings used in three types of scrap 
charging tests in an electric arc furnace, weight percent 





Batch 
1 


Batch 
2 




Batch 3 


Stamp- 




Large 


Small 


ings 


Al ... 


0.46 


0.83 


NA 


2.15 


0.046 


C... 


.41 


.56 


0.051 


1.66 


.67 


Cr... 


.16 


.40 


.050 


.47 


.49 


Cu... 


.18 


.81 


.20 


1.18 


.087 


Fe... 


. 98.5 


95.7 


99.5 


91.5 


97.2 


Mn .. 


.12 


.43 


.036 


.69 


.41 


Ni . . . 


.17 


.30 


.11 


.31 


.70 


P.... 


.008 


.034 


.011 


.070 


.013 


S... 


.030 


.045 


.037 


.048 


.008 


Si... 


<.01 


.85 


.011 


1.86 


.26 


Sn... 


NA 


.019 


.009 


.037 


.005 


Ti .. . 


NA 


<.01 


<.01 


<.01 


.003 



NA Not analyzed. 

1 Pb and Zn contents of the scrap samples were not analyzed since these 
constituents volatilized from the furnace during charge meltdown and were 
recovered in the dust product. 



For all tests, the furnace was preheated, and the initial charge 
of 450 lb of shredded auto scrap, metallurgical coke reductant, and 
slag formers (pebble lime and quartz) was topcharged to the fur- 
nace by means of a charging bucket. After the initial charge was 
melted, continuous feeding commenced. For the tests in which cold 
scrap was continuously charged, the feed material was fed into an 
opening in the side of the feed chute between the furnace and the 
preheater, as shown in figure 7. The cold scrap to be preheated 
was fed directly into the preheater shown in figure 8. Preheated 
scrap temperatures were calculated from data obtained by taking 
samples of scrap as they entered the furnace from the preheater 
and immediately immersing the samples into a measured amount 
of water. 

During the backcharging tests, only a portion of the initial 
450-lb charge was melted. At that point, the furnace roof was swung 
aside, and the first backcharge, consisting of 675 lb of shredded 
scrap, was loaded as shown in figure 9. When this was melted suffi- 
ciently to allow another 675 lb of shredded scrap to be added, the 
same procedure was followed. The remainder of the test was iden- 
tical to the continuously charged tests. 




/ Probe i'Cyclone J Flask ^ Particulate filter 5 Impingers 6' Thermometer 
7 Check valve 8 Connecting cord 9 Vacuum gage 10 Coarse adjust valve 
// Fine adjust valve 12 Oiler 13 Vacuum pump 14 Filter 15 Dry-gas meter 
16 Orifice tube 17 Incline manometer 18 Solenoid valves 19 Pilot <?C Thermo- 
couple 21 Temperature recorder 



Figure 6.— Schematic diagram of apparatus used for stack gas measurements and sampling. 



17 




Figure 7.— Continuous feeding of cold scrap into feed chute extending through roof of electric arc furnace. 



18 




Figure 8.— Continuous charging of preheated scrap into electric arc furnace. 



19 




Figure 9.— Backcharging of cold scrap into electric arc furnace. 



20 



RESULTS 
Continuously Charged Preheated Scrap 

Table 4 summarizes the averaged experimental data for 12 
tests made with preheated scrap. One standard deviation from the 
means is shown. The feed rates ranged from 22.9 to 66.2 Ib/min, 
and the power input ranged from 440 to 682 kW. At the feed rates 
used, no unmelted scrap buildup was noted. For the continuous feed- 
ing period, electrical energy consumption ranged from 592 to 810 
kWh/st of charge. Overall energy consumption for these tests 
ranged from 786 to 924 kWh/st of scrap. Test-to-test variations 
in scrap feed rate owing to the fragmented scrap's tendency to hang 
up in the charge bin primarily were responsible for the range in 
energy consumption. The temperatures of the furnace offgases 
introduced into the preheater ranged from 1,020° to 1,200° F. Exit 
gas temperatures usually were less than 390 ° F owing to transfer 
of heat to the preheated scrap and infusion of air into the feed inlet 
during the continuous charging operation. Preheated scrap temper- 
atures ranged from 840 ° to 1 , 1 10 ° F. Stack gas temperature ranged 
from 101 ° to 135° F, while the flows were 1,382 to 1,783 scfm. 
The stack gas contained 0.03 to 2.24 gr/dscf of particulate. Mois- 
ture contents of the gases ranged from 0.09 to 1.92 wt pet. 

Continuously Charged Cold Scrap 



Figure 10 shows the actual electrical energy consumed in the con- 
tinuous addition, meltdown, and refining periods of the two types 
of continuous scrap charging tests, along with tests for conventional 
backcharged scrap. It can readily be seen that both types of con- 
tinuous charging tests consumed more electrical energy than did 

Table 4.— Averaged experimental data for steelmaking tests 

using continuously charged preheated and cold scrap, and 

backcharged cold scrap, 

Continuously charged scrap Backcharged 
Preheated Cold cold scrap 

Total test time min . . 88+10 97 + 7 80 + 3 

Av power input kW . . 577 + 51 579±31 569±14 

Av scrap feed rate Ib/min . . 44 ± 1 38 ± 9 NAp 

Overall energy consumption 

kW-h/st scrap . . 829 ±43 88 ±51 637 ±23 
Electrode consumption 

Ib/st scrap .. 16±3 15±3 9±1 

Tap temperature °F . . 2,992±115 3,030 + 115 2,892±107 

Metal Ib/st scrap . . 1 ,856 ± 52 1 ,888 ±71 1 ,859 ± 95 

Slag Ib/st scrap . . 245 ± 66 241 ± 78 1 40 ± 1 1 

Stack dust Ib/st scrap . . 56±21 189±65 14±7 

Metal recovery pet . . 93 94 93 

NAp Not applicable. 

' Average of 4 tests. A suitable stack dust collection system was not availa- 
ble for the other 10 tests. 



The averaged experimental data for 14 tests conducted with 
continuously charged cold scrap also are shown in table 4. The scrap 
feedrate ranged from 30.5 to 63.7 Ib/min, and the power input 
ranged from 498 to 657 kW. Electrical energy consumption for the 
feeding period ranged from 630 to 846 kWh/st of charge. Over- 
all, the energy consumption was 818 to 1,084 kWh/st of charge, 
and the electrode use ranged from 10.5 to 19.1 Ib/st of scrap. Stack 
gas temperatures for 4 of the 14 tests ranged from 191 ° to 270° 
F, with flow rates of 1,220 to 1,377 scfm. Particulate concentra- 
tions ranged from 0.01 to 0.53 gr/dscf. Moisture contents of the 
gases ranged from 0.86 to 2.15 wt pet. 



Scrap Backcharging Tests 

The averaged experimental data for three scrap backcharge tests 
also are presented in table 4. In these tests, the power input ranged 
from 533 to 579 kW for the melt down of the 1 ,800-lb scrap charges. 
Electrical energy consumption ranged from 612 to 658 kWh/st of 
charge. Electrode consumption of 10.2 Ib/st of scrap or less was 
noted in these tests. Stack gases having temperatures of 126° to 
146° F flowed at a rate of 1,385 to 1,490 scfm. Particulate con- 
centrations ranged from 0.11 to 0.64 gr/dscf. The moisture con- 
tent of the gases was 1.31 to 1.75 wt pet. 

Comparisons of Melting Techniques 

The data indicate that continuously charged preheated scrap 
consumed 5.3 pet less electrical energy than when cold scrap was 
continuously charged during meltdown. Tests with continuously fed 
preheated and cold scrap charges overall consumed 829 and 888 
kW-h/st of scrap, respectively. Overall, the preheated scrap charges 
consumed 6.8 pet less electrical energy than did the cold-charged 
scrap. 



Cold scrap 



Preheated 
scrap 



Backcharged 
scrap 







eltdown period 



p^:^ Continuous 
^ additions 



Refining period 



400 800 

ELECTRICAL ENERGY, 

kW-h/st 

Figure 10.— Actual electrical energy consumption for three 
types of scrap charging tests. 



21 



the tests made with conventional scrap backcharges. Higher energy 
consumption may be attributed (1) to the inability to continuously 
charge at a rate fast enough to use all of the available energy, or 
(2) to less efficient heat transfer using the preheating system. The 
first hypothesis is supported by the open-bath conditions and the 
higher bath temperature observed during continuous charging opier- 
ations. No significant differences were noted in the quantity of metal 
produced in the three types of tests. 

It became apparent that the scrap feed rate was not balanced 
with the high power input used. The scrap charging and preheat- 
ing equipment was not large enough to feed the scrap at a rate cor- 
responding to the power input used. One suggested technique to 
address this problem is choke feeding, whereby the scrap feed rate 
is sufficient to allow an accumulation of scrap in the furnace to pre- 
vent open-bath conditions. The choke feeding technique used for 
prereduced iron ore pellets resulted in higher melt rates and 
increased productivity (26). Another technique to employ during 
continuous feeding is use of a foamy slag (27). The radiation of 
arc energy to the sidewalls is reduced since the tips of the elec- 
trodes are surrounded by scrap or a foamy slag. In the present case, 
the equipment and electric-arc furnace were not large enough to 
permit testing of either technique. 

Analytical data from backcharged tests indicated the furnace 
dust contained less zinc (32.5 wt pet) than the dusts derived from 
the continuous feeding charges. Both types of continuous-charging 
tests averaged approximately 100 lb more slag than did tests made 
with conventional scrap backcharges. This probably was because 
of dissolved refractory materials since open-bath conditions dur- 
ing the continuous feeding period caused a portion of the arc energy 
to be radiated to the furnace sidewall . In the case of conventional 



scrap backcharge practice, the furnace sidewall is shielded most 
of the time by the unmelted scrap. In the absence of an open bath, 
less of the arc energy is radiated to the sidewall. 

Some differences in stack gas temperatures and flows were 
observed among the three types of scrap charging tests, as shown 
in table 5. Stack gas temperature in the preheating tests was lower 
because heat was absorbed by the scrap fed into the preheater unit. 
The higher gas flow was attributed to the ingress of air into the 
scrap feed inlet of the preheater during the continuous addition 
period. Stack gas from the preheated scrap tests also contained the 
highest concentrations of particulate (0.97 gr/dscf). A portion of 
the organic material (such as upholstery), iron oxide scale, and other 
materials that mechanically separated from the continuously fed 
scrap as it tumbled in the rotating preheater became entrained as 
particulate in the furnace offgas stream. In the other type of test, 
the scrap was charged direcdy into the furnace, and part of the dirt, 
glass, etc., reported to the slag phase. 



Table 5. — Averaged data for stack gas measurements of the 
three types of scrap charging tests. 



Continuously charged scrap Backcharged 
Preheated Cold only 



Sampling duration min . 

Temperature of gas °F . 

Gas velocity ft/s . 

Flow rate scfm . 

Moisture content wt pet . 

Particulate cone gr/dscf . 



21.6±8.4 33.0±8.8 18.3±1.3 

120±11 220±35 135±10 

28±2 26.5±1.4 25.5±1.3 

1,615 ±132 1,302 ±73.4 1,433 ±53.2 

1.3±0,4 1.6±0.6 1.5±0.2 

0.97 ±0.74 0.24 ±0.24 0.29 ±0.3 



SUMMARY AND CONCLUSIONS 



Process offgases generated during oxygen blowing of a scrap- 
molten iron charge can be used effectively to preheat the next scrap 
charge in a '/4-st capacity BOF. A maximum of 28 pet unheated 
scrap can be used in normal BOF operations. When scrap is pre- 
heated to 1 , 150 ° F, scrap in the charge can be increased to a maxi- 
mum of 40 pet, representing a 43-pct increase in scrap utilization. 
Thermal energy recovered using 40 pet scrap represents 26 pet of 
the heat input required to melt the scrap. Scrap preheating in a BOF 
decreases energy consumption, allows the use of additional scrap 
to augment hot metal supplies when scrap is relatively cheap or 
the demand for hot metal exceeds the supply, and increases scrap 
utilization. 



In electric arc furnace steelmaking operations, continuously fed 
fragmented scrap preheated by furnace offgases consumed about 
5 pet less electrical energy for the meltdown and 7 pet less overall 
than did similarly charged cold scrap. Data indicated that the pre- 
heated and cold charges of fragmented scrap continuously fed into 
the furnace consumed more electrical energy and electrode materials 
than did conventional backcharged scrap. The higher energy con- 
sumption is not necessarily characteristic of continuous feeding but 
may be due to the way the testing was conducted. Significant differ- 
ences in stack gas temperatures, flows, and particulate concentra- 
tions were noted among the three types of scrap charging tests. Stack 
gases from scrap preheating tests exhibited the highest flow rate 
and the lowest temperature. 



22 



REFERENCES 



1. Schottman, F. J. Iron and Steel. Ch. in BuMines Minerals Yearbook 
1985, V. 1, pp. 553-571. 

2. Kotraba, N. L. Technology Alters Steelmaking. Am. Met. Mark,, 
V. 69, No. 237, Dec. 9, 1981, p. 16. 

3. Hale, R. W. Energy Use Patterns in Metallurgical and Nonmetallic 
Mineral Processing (Phase 8 — Opportunity To Improve Energy Efficiency 
in Production of High-Priority Commodities Without Major Process 
Changes) (contract S0144093, Battelle Columbus Laboratories). BuMines 
OFR-117(3)-76, 1975, 80 pp.; NTIS PB 261 152. 

4. Hunter, W. L., and J. E. Tress. Preheating Prereduced Briquettes 
and Development of Continuous Steelmaking. Paper in Proceedings of the 
27th Electric Furnace Conference. AIME, v. 27, 1970, pp. 122-125. 

5. Tress, J. E., W. L. Hunter, and W. A. Stickney. Continuous Charging 
and Preheating of Prereduced Iron Ore. BuMines RI 8004, 1975, 10 pp. 

6. Kobrin, C. L. Preheating Scrap for the BOF. Iron Age, v. 210, No. 
2, 1968, pp. 57-59. 

7. Moresco, A. J. Scrap Preheat Operation. Iron Steel Eng., v. 46, No. 
6, 1969, pp. 103-105. 

8. Laws, W. R. Prospects for Scrap Preheating for the Basic Oxygen 
Furnace. Steel Times, v. 200, No. 9, 1972, pp. 679-682. 

9. Kemner, W. F., The Operating, Economic and Quality Considera- 
tions of Scrap Preheating With the Basic Oxygen Process. Blast Fum. Steel 
Plant, V. 57, No. 12, 1969, pp. 1007-1012. 

10. Chatterjee, A. Economics of Preheated Scrap Usage in the LD Proc- 
ess. Iron Steel Int., Aug. 1973, pp. 325-331. 

11. Meschter, E. Preheaters in the Cold. Am. Met. Mark., v. 90, No. 
187, Sept. 27, 1982, Steelmaking Suppl., p. 6A. 

12. Kishida, T. A. Ukai, S. Sugiura, and S. Asano. Scrap Preheating 
by Exhaust Gas From Electric Arc Furnaces. Iron Steel Eng., v. 60, No. 
II, 1983, pp. 54-61. 

13. Industrial Heating. Scrap Preheating With Electric Arc Melting Fur- 
nace Off Gases Improves Operating Efficiency. V. 50, No. 7, 1983, p. 28. 

14. Watanabe, H., M. Iguchi, and T. Maki. Scrap Preheater for Elec- 
tric Arc Furnace. Iron Steel Eng., v. 60, No. 4, 1983, pp. 45-50. 



15. Tomizawa, F. and E. C. Howard. Arc Furnace Productivity in the 
1980's. Iron Steel Eng., v. 62, No. 5, 1985, pp. 34-37. 

16. Finkl, C. W. Benefits of Pre-Heating Scrap. J. Met., v. 17, No. 1, 
1965, pp. 67-70. 

17. Einerkjaer, J. Preheating of Electric Furnace Scrap, Iron Steel Eng., 
V. 47, No. 4, 1970, pp. 51-55. 

18. Rudzki, E. M., R. J. Reinbold, and B. K. Pease. Scrap Preheating 
in an Electric Melt Shop. J. Met., v. 25, No. 2, 1973, pp. 38-43. 

19. Remalia, D. L. Scrap Preheating for Increased Productivity. Paper 
in 37th Electric Furnace Conference Proceedings. AIME, v. 37, 1979, pp. 
315-319. 

20. Mahony, H. A. Rapid Steel Scrap Melting With Gas-Oxygen Burners. 
U.S. Pat. 3,234,010, Feb. 8, 1966. 

21. . Smelting of Fine Iron Ores to Molten Iron in Induc- 
tion Furnaces. U.S. Pat. 3,235,374, Feb. 15, 1966. 

22. Tomizawa, F., and E. C. Howard. Scrap Preheating Within a Clean 
House Enclosure and Associated Operation Benefits. Paper in 42d Electric 
Furnace Conference Proceedings. AIME, v. 42, 1984, pp. 79-91. 

23. Drost, J. J., C. B. Daellenbach, W. M. Mahan, and W. C. Hill. 
Thermal Energy Recovery by Basic Oxygen Furnace Offgas Preheating of 
Scrap. BuMines RI 7929, 1974, 8 pp. 

24. Mahan, W. M., C. B. Daellenbach. Thermal Energy Recovery by 
Basic Oxygen Furnace Offgas Preheating of Scrap. Paper in Procedings 
of Symposium on Efficient Use of Fuel in the Metallurgical Industries. Inst. 
Gas Technology, IIT Center, Chicago, IL, Dec. 1974, pp. 457-465. 

25. Elger, G. W., R. H. Nafziger, J. E. Tress, and A. D. Hartman. 
Utilization of Scrap Preheating and Substitute Slag Conditioners for Elec- 
tric Arc Furnace Steelmaking. BuMines RI 9130, 1987, 26 pp. 

26. Nafziger, R. H., J. E. Tress, and W. L. Hunter. Rapid Addition 
of Charge Materials in Continuous Electric Furnace Steelmaking. Iron Steel- 
maker, v. 2, No, 5, 1975, pp. 33-37 

27. Caine, K. E., Jr. A Review of New Electric Arc Furnace Technolo- 
gies. Iron Steel Eng., v. 60, No. 10, 1983, pp. 45-47. 



23 



FLUORSPAR SUBSTITUTES IN STEELMAKING 



By R. H. NafzigeM and G. W. Elger^ 



ABSTRACT 



One goal of Bureau of Mines research is to establish the potential of abundant domestic resources 
as substitutes for imported materials. Fluorspar, which is used as a slag conditioner in steelmak- 
ing, is one of these materials. The Bureau has conducted research on substitutes for fluorspar such 
as colemanite (Cai^eOir^i^iO)^ fused boric acid, synthetic fluorspar, and used aluminum smelter 
potlining in basic oxygen furnace (BOF) operations, and on synthetic fluorspar, boric acid, 
hydroboracite (CaMgB60,,-6H20), used aluminum potlining and anode tailing wastes, and Sorel- 
flux B (ilmenite) in electric arc furnace steelmaking. For BOF slags, colemanite and fused boric 
acids were superior fluidizers to fluorspar when compared on the basis of boron versus fluorine 
concentration and were more stable. Synthetic fluorspar and used aluminum potlining provided 
equivalent slag fluidity to that noted by natural fluorspar. In electric arc furnace steelmaking, the 
boron-containing conditioners (hydroboracite and boric oxide) fluidized the slags better than those 
containing fluorine (synthetic fluorspar, used aluminum podining, and anode tailing wastes) and 
titanium (Sorelflux B), in that order. The results indicate that substitute fluidizers do not adversely 
affect the quality of the steel produced. 



INTRODUCTION 



Fluorspar is a critical mineral. Approximately 81 pet (553,000 
St) of domestic fluorspar consumption was imported in 1985 (1).^ 
Estimates for 1986, based on 9 months of data, show 94 pet of the 
domestic needs for fluorspar was imported (2). Approximately 33 
pet of the fluorspar consumed domestically is used as a fluidizer 
in steelmaking. Most steelmaking operations require a slag fluidizer 
to promote the required reactions, to increase productivity, to 
improve tapping operations, and to conserve energy. Fluorspar, 
as an auxiliary flux, promotes the rapid solubility of lime and rapid 
slag formation, in both the BOF and electric arc furnace. This also 
improves BOF steelmaking operations by lowering frothing and 
sparking during the oxygen blowing period. Over 181,000 st of 
fluorspar was consumed in steelmaking in 1985 (1). In 1985, BOF's 
used approximately 2.7 lb of fluorspar per short ton of raw steel 
produced (1). The electric arc furnace consumes approximately 2. 1 
lb of fluorspar per short ton of raw steel . 

If a satisfactory substitute can be found, U.S. dependence on 
foreign sources of suitable fluorspar would diminish and import 
costs could decrease. A number of potential fluorspar substitutes 
that are more readily available from domestic sources than fluor- 



' Research supervisor. 
^ Research chemist. 

Albany Research Center, Bureau of Mines, Albany, OR, 
' Italic numbers in parentheses refer to items in the list of references at the end of 
this paper. 



spar are not being used. Several of these potential substitutes are 
wastes of which the disposal presents environmental problems. 

The search for fluorspar substitutes as a flux conditioner began 
a number of years ago with work conducted at Stelco (Steel Com- 
pany of Canada). Results indicated that calcium borates and man- 
ganese ore were similar to fluorspar in their ability to form a basic 
slag rapidly in BOF operations. Tests in a 500-st capacity open- 
hearth furnace confirmed these observations (3). QIT-Fer et Titane, 
Inc., has been promoting its Sorelflux (ilmenite) as a good fluor- 
spar substitute for several years. Most recently, trials were con- 
ducted in a 4-st capacity electric arc furnace using spent aluminum 
potlining material . All of the heats provided steel within specifica- 
tions. From a qualitative standpoint, the slags with the podining 
fluidizer were more fluid than those obtained when fluorspar was 
used (4). 

The objective of the research reported in this review paper is 
to evaluate the use of readily available domestic materials or wastes 
as substitutes for fluorspar slag conditioners in BOF and electric 
arc furnace steelmaking. Substitute materials were compared with 
fluorspar on the basis of both viscometry measurements and visual 
observations made of the molten bath before and after the fluidizer 
addition. This investigation is part of the Bureau program to develop 
technology that emphasizes the reuse of recycled materials and to 
help meet the goal of substituting abundant domestic materials for 
imported critical materials. 



24 



SUBSTITUTE MATERIALS 



For BOF steelmaking, colemanite (Ca2B60n-5H20), fused 
boric acid, synthetic fluorspar made from fluosilicic acid at the 
Bureau's Albany (OR) Research Center, and used aluminum smelter 
podining materials were used as conditioning substitutes and were 
compared with reagent- and ceramic-grade fluorspar. Both the 
colemanite and the fused boric acid were obtained commercially. 
The synthetic fluorspar was prepared from fluosilicic acid, a 
byproduct of the acidulation of fluorapatite (Caio(P04)6F2), 
produced in the manufacture of fertilizers (5). The used aluminum 
smelter potlining material was provided by Alcoa, Pittsburgh, PA. 
Analyses of these materials are summarized in table 1 . 

Table 1 . — Average chemical composition of slag 

conditioners used in BOF steelmaking tests (7, 10), weight 

percent 



Fluorspar 


Colemanite 'bo'? ^ynthetic 
acid ^'^°'^P^' 


Used 


Reagent- Ceramic- 
grade grade 


potlining 
Lump! Pellet^ 



Al.. 
B.. 
C. 
Ca. 
F. . 
Mg. 
Na. 
P.. 
S.. 
Si.. 



ND 
NDt 

ND 
49.4 
48.0 
NDt 
NDt 

ND 

ND 
<.05 



0.01 

NDt 

.38 

49.9 

45,0 

.12 

.04 

.12 

.07 

<.05 



ND 
12.2 

ND 
18.3 
NDt 
1.4 
<.04 
ND 
ND 
3.0 



ND 
28.2 
ND 
NDt 
NDt 
NDt 
.01 
ND 
ND 
NDt 



ND 

0.3 

2.2 

39.0 

31.7 

.3 

.2 

1.3 

.8 

3.2 



6.9 
ND 
34.8 

2.1 
12.5 
.06 

7.4 
ND 
.2 

2.5 



6.7 

ND 

38.1 

1.9 

11.5 

.06 
6.6 
ND 
.2 
2.9 



ND Not determined. NDT 

1 Minus % in plus 10 mesh. 

2 Minus % plus V2 in. 



Not detected. 



In the electric arc furnace steelmaking tests, three types of slag 
conditioners were used: fluorine-, titanium-, and boron-containing 
compounds or materials. The fluorine group included natural fluor- 
spar, synthetic fluorspar, used aluminum smelter potlining, and 
anode tailing wastes. The natural fluorspar was purchased from a 
commercial supplier and was the type normally used in steelmak- 
ing operations. The synthetic fluorspar was similar to that used in 
the BOF tests. The used potlining and anode or tailings were 



obtained from Kaiser Aluminum and Chemical Co., Spokane, WA. 
Both were from aluminum reduction cells. The linings of alumi- 
num reduction cells must be replaced regularly. The linings are car- 
bonaceous and become impregnated with fluorine and sodium 
compounds while in service. The material accumulates at a rate 
of approximately 190,000 st/yr and presents a disposal problem 
for the aluminum industry (6). The anode tailing wastes were the 
ends of the carbon anodes. 

The titanium-containing material was Sorelflux B. This was 
furnished by QIT-Fer et Titane, Inc., and was mostly ilmenite and 
feldspar, with a small amount of hematite. 

The boron-containing conditioners included boric oxide and 
Gerstley borate. Boric oxide was prepared by fusing and grinding 
boric acid. The Gerstley borate, which consisted of hydroboracite 
(CaMgBftO,, -61120), is used as a glaze in the ceramics industry. 
The Gerstley borate was heated to 800 ° F to drive off the contained 
water. The chemical analyses of all these slag conditioners are shown 
in table 2. 



Table 2. — Chemical composition of slag conditioners used 
in electric arc furnace steelmaking tests (12), weight percent 



Fluorspar 



Used 



Butt Sorel- Boric Gerstley 



Natural Synthetic potlining tailings flux B oxide borate 



Al.. 

B. . 

C. . 
Ca. 
F.. 
Fe. 
Mg. 
Mn. 
Na. 
P. . 
Pb. 
S.. 
Si.. 
Ti.. 

v.. 



2.11 
ND 

1.27 
40.6 
34.8 
.28 
.27 
ND 
ND 
ND 

1.0 
<.39 

3.70 
ND 
ND 



0.09 

ND 

.71 

45.3 

40.5 

ND 

.25 

ND 

ND 

ND 

.34 

.19 

1.77 

ND 

ND 



11.3 
ND 
1.86 
2.10 

21.5 
ND 
.10 
ND 

25.0 
ND 
ND 
.27 
.81 
ND 
ND 



2.02 
ND 
63.8 
.53 

7.5 
ND 
.04 
ND 

4.1 
ND 
ND 

2.85 
.31 
ND 
ND 



1.97 
ND 
ND 
.66 
ND 
38.8 
1.79 
.12 
ND 
.021 
ND 
ND 
2.65 
20.3 
.25 



ND 
22.2 
ND 
ND 
ND 
ND 
ND 
ND 
ND 
ND 
ND 
ND 
ND 
ND 
ND 



0.50 

9.72 

ND 

12.7 

ND 

.23 

2.01 
ND 

3.75 
ND 
ND 
ND 

4.10 
ND 
ND 



ND Not determined. 



BOF EXPERIMENTS 



LABORATORY-SCALE VISCOMETRY TESTS 

Laboratory-scale viscometry experiments were conducted on 
boron-containing slags and results were compared with fluorspar 
determinations at the Bureau's Twin Cities (MN) Research Center 
(7). Two master slags were prepared in the BOF and ground to 
minus 100 mesh. Preheated lime was added to the molten slag sam- 
ples in the viscometer to adjust the basicity or CaO-Si02 ratio 
(usually more than 3:1). The viscometer was of the rotating- 
concentric cylinder type and has been described in detaU by Kilau 
(7). The wire-wound resistance furnace and viscosity transmitter 
also have been described (7). The addition of a pressure transducer, 
X-Y recorder, and temperature programmer allowed continuous 
viscosity measurements. 

Kilau (7) found that when 2.2 to 3. 1 wt pet fluorine was added 
as reagent- or ceramic-grade fluorspar, the viscosity increased with 



increasing slag basicity. At approximately 3 wt pet fluorine, a 
fluidizing limit was suggested, above which foaming and crusting 
may occur. Based on analyses of quenched slags, from 5 to as much 
as 26 pet fluorine was lost at temperatures up to approximately 
1 ,500 ° C . On the basis of elemental boron and fluorine concentra- 
tions, boron added as colemanite is a superior fluidizer to fluor- 
spar for slags with a basicity of 3. This is illustrated in figure 1 
(7). Fused boric acid proved to be slighfly superior to colemanite 
as a slag fluidizer. Both colemanite and fused boric acid showed 
no losses of boron upon heating. No fluidizing limit for fused boric 
acid was noted by up to 4.4 wt pet boron levels. 

For synthetic fluorspars, up to 58 pet fluorine loss was noted 
at 1,000° C by Kilau (8). However, it was shown that improved 
synthetic fluorspars only lost up to 14 pet fluorine (8). The improved 
synthetic fluorspar used limestone in the precipitation of CaF2 from 
the fluosilicic acid. Also, pH control was improved. Such synthetic 



25 



fluorspars can differ greatly in tiieir stability and ability to fluidize 
BOF slags, depending upon their method of preparation. As shown 
in figure 2 (8), the improved synthetic fluorspars are comparable 
to natural (ceramic-grade) fluorspar. However, Kilau showed that 
a commercially prepared synthetic fluorspar was somewhat inferior 
to the natural fluorspar (fig. 2). 

TESTS IN THE BOF 
Equipment and Procedures 

Experiments to evaluate synthetic fluorspars and used alumi- 
num smelting podining materials were conducted in the Bureau's 
500-lb-capacity BOF at the Twin Cities (MN) Research Center. 




1,150 1,200 



1,250 1,300 1,350 

TEMPERATURE, °C 



1,400 1,450 



Figure 1. — Viscosity-temperature profiles for fluorspar, 
colemanite, and fused boric acid additions to BOF slags (7). A, 
2.0 pet B from added dehydrated colemanite (13.6 pet); B, 1.9 
pet B from added fused boric acid (6.7 pet); C, 2.9 pet F from 
added ceramic-grade fluorspar (6.4 pet); D, 3.1 pet B from added 
fused boric acid (1 1 .0 pet). 



This furnace has been described previously by Drost (9). Two syn- 
thetic fluorspars (the improved and commercial varieties) were com- 
pared with natural ceramic-grade fluorspar using identical 300-lb 
charges of hot metal and 80 lb of shredded automotive scrap (10-11). 
The synthetic fluorspars were briquetted to decrease dusting losses 
during BOF ojjerations. Fluidizers were added on an equivalent fluo- 
rine basis (2.0-2.6 lb). The charges also consisted of 30 lb lime. 
All heats were blocked in the furnace using 2 lb of ferrosilicon and 
1.5 lb of electrolytic manganese. Blocking deoxidizes the heat to 
maintain a constant carbon content. Each fluorspar test consisted 
of three heats. The BOF was blown at 25 scfm oxygen for 10 to 
12 min. Offgases were passed through a wet scrubber. The scrub- 
ber water and solids were collected and submitted for analyses, along 
with the metal and slag samples. Additional details were described 
by Spironello (10-11). 

The same furnace was used to evaluate used aluminum smelter 
potlining material as a substitute fluidizer. Shredded automotive 
scrap (90 lb) was added to the furnace, followed by hot metal 
(310 + 15 lb). The oxygen flow was set at about 24 cfm. Approxi- 
mately 1 min after ignition, 30 lb lime was charged, followed by 
the potlining material (4 lb). Ferrosilicon and silicomanganese were 
added after the oxygen blow to block the heats (10-11). 

Results 

When synthetic fluorspar slags were used, the furnace oper- 
ated satisfactorily. Metal and slag analyses are shown in table 3 
(8). Kilau found that the steel produced was satisfactory for all of 
the fluorspars tested, despite relatively high phosphorus levels in 
the synthetic fluorspar (table 1). MgO concentrations in the slag 
(table 3) suggested that no excessive refractory consumption resulted 
from the use of the synthetic fluorspars (8). The resulting slags were 
subjected to laboratory-scale testing as described previously in this 
paper. Results are shown in figure 3 (8). It can be seen that the 
commercial synthetic fluorspar appears greatly inferior to either 
the improved synthetic fluorspar or to the natural fluorspars. 



100 
90 
80 

70 

cl 
^ 60 

3~ 

>- 50 

CO 

O 40 

o 

tn 

> 30 
20 



~i 1 \ r 

More fluid slags 



1 I I r 

Less fluid slogs 



Low 



Apporent 
solidification 



High 



1,300 1,340 




1,380 1,420 

TEMPERATURE, 



1,500 



1,540 



Figure 2.— Viscosity-temperature profiles for natural and syn- 
thetic fluorspar additions to BOF slags (8), A,, Natural fluorspar 
(ceramic grade), 2.7 pet F; A2, natural fluorspar (ceramic grade), 
2.2 pet F; B,, synthetic fluorspar (improved product), 3.2 pet F; 
B2, synthetic fluorspar (improved product), 2.9 pet F; C, synthetic 
fluorspar (commercial product), 2.9 pet F; Di, synthetic fluor- 
spar, 2.7 pet F; D2, synthetic fluorspar, 1.9 pet F. 



100 
90 
80 
70 

Q- 

^ 60 
p- 

>- 50 

<n 

o 40 
(J 
</) 

> 30 
20 



10 - 



"I i 1 1 r 

More fluid slogs 



"I 1 1 r 

Less fluid slogs 



Low 



Apparent 
solidification 
temperature 



-High 



1,280 1,320 




1,360 1,400 

TEMPERATURE, 



1,480 1,520 



Figure 3. — Viscosity-temperature profiles for pilot-scale BOF 
steelmaking slags to which natural and synthetic fluorspars were 
added (8). Ai, Overblown BOF slag with natural fluorspar, 0.7 
pet F, 35.9 pet total Fe, 3.6 basicity; A2, normal practice BOF 
slag with natural fluorspar, 1.4 pet F, 19.7 pet total Fe, 3.8 basic- 
ity; B, BOF slag with synthetic fluorspar (improved product) 1 .3 
pet F, 17.6 pet total Fe, 3.3 basicity; C, BOF slag with commer- 
cial synthetic fluorspar, 1 .4 pet F, 13.3 pet total Fe, 4.4 basicity. 



26 



Table 3.— Chemical analyses of BOF slags 
and metals (8), weight percent 



Table 4.— Steel analyses from evaluation of used potlining 
in a BOF (10), weight percent 





Natural 


Bureau 


snythetic 


Commercial syn- 




fluorspar 


fluorspar 


thetic fluorspar 




Metal 


Slag 


Metal 


Slag 


Metal 


Slag 


AI203 .... 


NAp 


0.35 


NAp 


0.35 


NAp 


0.36 


c 


0.28 


NAp 


0.05 


NAp 


0.07 


NAp 


CAO 


NAp 


42.1 


NAp 


50.5 


NAp 


53.2 


F 


NAp 


1.5 


NAp 


1.3 


NAp 


1.2 


Fe2+ .... 


NAp 


16.8 


NAp 


12.6 


NAp 


13.3 


MgO .... 


NAp 


4.9 


NAp 


2.6 


NAp 


2.7 


Mn 


.23 


6.1 


.26 


4.6 


.29 


5.5 


P 


<.01 


.28 


<.01 


.36 


<.01 


.51 


S 


.011 


.096 


.014 


.076 


.016 


.100 


Si 


.12 


NAp 


.24 


NAp 


.12 


NAp 


SiOa 


NAp 


12.8 


NAp 


15.1 


NAp 


14.8 


Total Fe . . 


NAp 


20.7 


NAp 


16.5 


NAp 


16.0 



NAp Not applicable. 

The results of Kilau (8) also indicated that a fluidizer substi- 
tute need only be effective in the early stages of a blow to solubi- 
lize the dicalcium silicate. Toward the end of the blow, increasing 
iron contents of the slag could serve to provide the required fluidity. 
Hence, a relatively unstable synthetic fluorspar could be effective 
in BOF steelmaking. In some cases, a BOF could be operated with- 
out fluorspar, especially for low-carbon steels in which the longer 
blowing time could permit adequate refining and produce sufficient 
iron oxide in the slag to effect satisfactory fluidization. 

Chemical analyses of the steel produced in the BOF using alu- 
minum smelter podining material and fluorspar are shown in table 
4 (10). Spironello found that higher sulfur levels resulted from the 
use of potlining material because the sulfur content was higher in 
the hot metal charge. All slags were fluid and permitted satisfac- 
tory sampling and tapping. Slag analyses are presented in table 5, 
also from Spironello (10). Based on the MgO contents of the slags, 
there were no significant differences in refractory lining attack when 
potlining was used. The higher AI2O3 and Na20 contents resulted 
from the aluminum and sodium in the potlining material (10). 



Heat 



Mn 



Fluorspar: 

1 

2 

Potlining 
lump:' 

1 

2 

Potlining 
pellets:2 

1 

2 



0.46 
.47 



.35 
.57 



.42 
.33 



<0.10 
<.10 



<.10 
<.10 



.10 
.10 



0.42 
.45 



.36 
.34 



.46 
.58 



0.006 
.013 



.010 
.010 



.010 
.010 



0.008 
.009 



.019 
.015 



.023 
.017 



0.10 
.15 



.10 
.10 



.10 
.18 



1 Minus % in plus 10 mesh. 

2 Minus % plus V2 in. 



Table 5. — Slag analyses from evaluation of used potlining 
in a BOF (10), weight percent 



Heat AI2O3 CaO F Fe^* MgO Mn NazO P 



SiO: 



Total 
Fe 



Fluorspar: 

1 0.44 53.6 1.4 12.4 5.4 4.3 <0. 02 0.55 0.063 13.5 17.3 

2 73 56.4 1.8 11.0 5.0 5.0 <.02 .69 .072 17.5 11.1 

Potlining 
lump:' 

1 2.8 50.5 .91 11.8 4.7 5.5 .95 .47 .087 14.8 16.2 

2 3.2 53.1 .85 14.5 4.5 5.4 .95 .52 .084 13.2 17.5 

Potlining 
pellets:^ 

1 1.7 56.6 .67 8.6 4.7 4.6 .47 .56 .100 15.1 11.6 

2 1.7 53.9 .65 11.6 4.8 5.1 .50 .57 .110 14.2 14.6 

1 Minus % in plus 10 mesh. 

2 Minus % plus V2 in. 

The product steels were hot- rolled and mechanically tested. 
Spironello found that the results showed yield and tensile strengths 
comparable to conventionally produced steels (10). 



ELECTRIC ARC FURNACE EXPERIMENTS 



EQUIPMENT AND PROCEDURES 

All tests were conducted in a 1-st-capacity, three-phase ac elec- 
tric arc furnace at the Bureau's Albany (OR) Research Center. 
Shredded automotive scrap, lime, and quartz were used as charge 
materials. After the feed had entered the furnace and the bath was 
molten, the slag conditioner was added. After about 5 min of heat- 
ing, the slag was removed from the bath. Then, 10 lb of silicoman- 
ganese was added, the bath was adjusted to the proper temperature, 
and the furnace was tapped. Both metal and slag samples were taken 
after slag removal and at the tap (12). 

RESULTS 

Of the three types of slag conditioners evaluated on the basis 
of visual observations (e.g., stickiness of the slag and ease of slag- 
metal separation), the Gerstley borate and B2O3 group appeared 
the most effective in increasing the fluidity of the individual slags. 



The resultant slags contained up to 0.5 wt pet B, while the tapped 
metal products contained 0.001 to 0.002 wt pet B. Boron at these 
levels is not considered a harmful constituent in most low- or 
medium-carbon steels. The constituent markedly increases the 
hardenability of steels at higher levels approaching 0.007 wt pet. 
The fluorine-containing materials were also effective in increas- 
ing slag fluidity. In the tests made with natural and synthetic fluor- 
spar slag conditioners, the tapped metal products contained between 
0.03 and 0.05 wt pet P. The synthetic fluorspar made from fluosi- 
licic acid obtained from a phosphate plant was not a source of phos- 
phorus contamination in the tapped metal products. The sodium 
constituent in aluminum potlining waste presented a slight fuming 
problem during scrap meltdown. In the third group, Sorelflux B 
was the least effective under the conditions used. Only visual com- 
parisons can be given since slag viscosity data were not obtained. 
Moreover, the slag compositions varied from test to test owing to 
erosion of furnace lining material. The furnace refractories suffered 
considerable damage owing to open-bath conditions (12). 



27 



SUMMARY AND CONCLUSIONS 



Experiments have shown that boron-containing compounds such 
as colemanite and fused boric acid fluidize BOF slags better than 
fluorspar on the basis of laboratory-scale viscosity determinations. 
The comparisons were made on equivalent amounts of boron and 
fluorine concentrations in the slags. Kilau demonstrated that the 
boron-containing slags showed no boron losses, whereas slags con- 
taining fluorspar were more unstable and lost fluorine by volatili- 
zation, resulting in increased slag viscosities and higher apparent 
solidification tempjeratures {7). Increased slag basicity increased the 
fluorine-containing BOF slag viscosity {7). The stabilities and 
fluidizing abilities of synthetic fluorspars in BOF slags varied con- 
siderably, depending upon their method of preparation, as shown 
by Kilau (8). Improved preparation methods {8) yielded synthetic 
fluorspars that compared favorably with natural fluorspar with 
respect to stability and fluidizing ability. On the other hand, a com- 
mercially prepared synthetic fluorspar was slightly inferior to nat- 
ural fluorspar (8). 

Tests in a pilot-scale BOF demonstrated that synthetic fluor- 
spar can be used as a substitute for natural fluorspar as a flux 



fluidizer without decreasing the quality of the metal produced. No 
excessive refractory wear was evident. Fluorspar fluidizers may 
need to be added only in the early stages of a blow since increasing 
iron oxide contents of the slag as the blow progresses in a BOF 
may provide the required fluidity. Low-carbon steels may require 
no fluidizer. Use of aluminum smelter potlining material as a sub- 
stitute fluidizer resulted in fluid slags and acceptable metal qual- 
ity. Refractory wear was nominal. 

Three groups of slag conditioners that contained boron 
(hydroboracite and fused boric acid), fluorine (natural and synthetic 
fluorspar and aluminum potlining and butt tailings), or titanium 
(Sorelflux B) were used as substitutes for natural fluorspar in elec- 
tric arc furnace steelmaking. The boron- and fluorine-containing 
additives were effective in increasing the fluidity of the slags, with 
the boron-containing materials more effective. Use of the boron- 
containing conditioners did not increase the boron levels in the melt- 
down metal. The tapped metal products contained 0.001 to 0.002 
wt pet B. Open-bath conditions caused refractory lining erosion. 



REFERENCES 



1. Pelham, L. Fluorspar. BuMines Minerals Yearbook 1985, v. I, 1987, 
pp. 419-428. 

2. . Fluorspar. Sec in BuMines Mineral Commodity 

Summaries 1987, pp. 52-53. 

3. Buxton, F. M., and P. A. Sandaluk. Fluorspar Substitutes in Steel- 
making. Ind. Heat., v. 40, 1973, pp. 288, 290, 292. 

4. Balding, P. C, D. R. Augood, and R. J. Schlager. Making Steel 
Using Spent Potlining Flux. Trans. Am. Foundrymen's Soc, v. 91, 1983, 
pp. 493-498. 

5. Nash, B. D., and H. E. Blake, Jr. Fluorine Recovery From Phos- 
phate Rock Concentrates. BuMines RI 8205, 1977, 16 pp. 

6. Balgord, W. D. Recycle of Potlining in the Primary Aluminum Indus- 
try. Opportunities for Technical Improvements. Paper in Proceedings of 
the Sixth Mineral Waste Utilization Symposium. IIT Res. Inst., Chicago, 
IL, 1978, pp. 324-333. 

7. Kilau, H. W., V. R. Spironello, and W. M. Mahan. Viscosity of 
BOF Slags Fluidized With Fluorspar, Colemanite, and Fused Boric Acid. 
BuMines RI 8292, 1978, 25 pp. 



8. Kilau, H. W., V. R. Spironello, I. D. Shah, and W. M. Mahan. 
Evaluation of Synthetic Fluorspar in BOF Slags. BuMines RI 8558, 1981, 
28 pp. 

9. Drost, J. J., C. B. Daellenbach, W. M. Mahan, and W. C. Hill. 
Thermal Energy Recovery by Basic Oxygen Furnace Offgas Preheating of 
Scrap. BuMines RI 7929, 1974, 8 pp. 

10. Spironello, V. R., and I. D. Shah. An Evaluation of Used Alumi- 
num Smelter Potlining as a Substitute for Flurospar in Basic Oxygen Steel- 
making. BuMines RI 8699, 1982, 11 pp. 

11. Spironello, V. R. An Evaluation of Aluminum Smelter Potlining as 
a Substitute for Fluorspar in Cupola Ironmelting and in Basic Oxygen Steel- 
making. BuMines RI 8775, 1983, 18 pp. 

12. Elger, G. W., R. H. Nafziger, J. E. Tress, and A. D. Hartman. 
Utilization of Scrap Preheating and Substitute Slag Conditioners for Elec- 
tric Arc Furnace Steelmaking. BuMines RI 9130, 1987, 26 pp. 



28 



RESEARCH ON BASIC STEELMAKING REFRACTORIES 



By T. A. Clancy^ and J. P. Bennett^ 



ABSTRACT 

The Bureau of Mines conducted research studies over a 10-yr period on basic refractories com- 
monly used in steelmaking processes. These studies dealt with properties of refractory raw materials 
(magnesia, dolomite, and natural flake graphite) and formulations containing these materials. High- 
temperature properties are reported for periclase (MgO) produced from seawater, brines, and magne- 
site (MgCOs). Improvements in high-temperature strength and slag resistance prop)erties are described 
for refractories produced from periclase grain altered by various additions. The properties of 14 
calcined domestic dolomites are described. The role of natural flake graphite in dolomite-carbon 
refractories and the feasibility of substitution by synthetic carbons are discussed. 



INTRODUCTION 



The Bureau of Mines has conducted extensive research on basic 
refractories commonly used in steelmaking processes. These labora- 
tory studies have focused on raw materials (periclase, dolomite, 
and natural flake graphite) and refractory formulations (magnesia 
and dolomite-carbon). Most of these studies have been directed at 
conserving natural resources via substitution or by improved 
materials performance. In the case of the refractory raw materials, 
laboratory studies generally consisted of chemcal, physical, and 
mineralogical characterization of materials from domestic sources. 
Refractory mix formulations, however, involved studies of high- 
temperature physical properties of new or modified mix formula- 
tions. High-temperature tests such as flexural strength, deforma- 
tion under load, and slag resistance evaluations are the most useful 
for comparing the high-temperature properties of various refrac- 
tories. 

Basic refractories are the preferred material in most steelmak- 
ing operations such as the basic oxygen furnace (BOF), the elec- 
tric arc furnace (EAF), the argon-oxygen decarburizing process 
(AOD), and ladles. As mentioned in a 1983 paper by Van Dreser 
and Neely (1),^ the BOF produces the greatest tonnage of steel in 
the United States, as in most of the world. Much attention has been 
focused on refractories used in this furnace, extending service life 
of refractory linings from 200 to 300 heats of earlier days to an 
average of 1,500 heats. The types of materials used with current 
BOF practice are shown in table 1 . Magnesite and periclase in the 
table distinguish only grades. Over 80 pet of the world's produc- 
tion of steel is produced in BOF's and EAF's (2). The EAF is con- 
tinually undergoing changes in steelmaking practices, bringing about 
changes in refractory requirements. These include changes such 



' Ceramic engineer, Tuscaloosa Research Center, Bureau of Mines, Tuscaloosa, 
AL. This paper is based upon work done under an agreement between the Univer- 
sity of Alabama and the Bureau of Mines. 

^ Italic numbers in parentheses refer to items in the list of references at the end 
of this paper. 



as scrap preheating, water-cooled panels for sidewalls and roofs, 
and EAF use as a feed source for AOD vessels. Material usage 
on EAF sidewalls and slaglines are shown in table 1 , roof refrac- 
tories, being mainly alumina-based materials, are not included. 
With the advent of water-cooled panels, the EAF is rapidly 
becoming a diminishing market for refractories. There is, however, 
a new market in gun mixes for water-cooled sidewall panels as main- 
tenance mixes, at slaglines and bottoms. At the slagline, the pre- 
dominant usage is magnesia-carbon (MgO-C) brick. Shop option 
and performance generally determine the type of MgO-C brick 
(amorphous graphite or flake graphite) and the carbon level (8-20 
pet residual carbon) to be used. As in the BOF, the principal mode 
of wear is decarbonization of the matrix. Consequendy, the major 



Table 1 .—Basic brick usage In steelmal<ing 

Application area Refractory 

Basic oxygen furnace: 

Bottom, cone Tempered, tar-bonded magnesite. 

Barrel Tempered, tar-bonded periclase. 

Trunnion pads MgO-carbon. 

Charge pad Burned, tar-impregnated MgO. 

Electric-arc furnace: 

Sidewall Direct-bonded periclase-chrome. 

Slagline Rebonded fused periclase-chrome. 

Hot spots MgO-C (flake graphite). 

Argon-oxygen decarburizer: 

Barrel, bottom Periclase-chrome or fused 

periclase-chrome grain. 

Tuyere pad Fused MgO-CraOa grain or 

dolomite. 
Ladle: 

Slagline Direct-bonded MgO-Cr203 or tar- 
bonded dolomite. 

Sidewall Chem-bonded MgO-Cr203 or tar- 
bonded dolomite. 

Impact pad Direct-bonded I\/Ig0-Cr203. 



29 



thrust in improving MgO-C brick is directed toward providing oxi- 
dation resistance or protection for the matrix. 

The AOD furnace is one of the most severe, highly corrosive 
environments to which refractories are exposed. Lining hfe has been 
extended from 20 to 30 heats in the early 1970's to in excess of 
100 heats today. The AOD furnace is used in conjunction with the 
EAF and is the principal means of production of stainless steel. 
Refractory usage associated with the AOD is shown in table 1 . 



The last major area of basic refractory usage is the ladle. New 
refractory demands evolve for ladle refractories as steel processes 
change to continuous casting and ladle refining. Commonly used 
basic linings are shown in table 1. 

This paper summarizes the research done at the Bureau's 
Tuscaloosa (AL) Research Center on the refractory raw materials, 
periclase, dolomite, and natural flake graphite used in basic 
refractories. 



RAW MATERIALS 



PERICLASE 

Conversion of steelmaking from the open hearth and Bessemer 
processes to the BOF and EAF, along with the shift to continuous 
casting, has resulted in the exposure of refractories to the much 
more hostile conditions associated with these newer processes. For 
example, the operating temperature of the newer metallurgical fur- 
naces has risen, with considerable changes in the requirements for 
refractory linings. Recent trends point toward the increased use of 
higher purity magnesia linings, which have become the most impor- 
tant refractory lining material in steelmaking. 

In the early 1950's, basic magnesia plants began producing 
high-purity-periclase (MgO) material from seawater and well and 
lake brines that competed economically with periclase from dead- 
burned magnesite (MgCOs) ores. Today both "natural" (magne- 
site) and "synthetic" (seawater, brines) magnesias ranging from 
93 to 99 pet MgO are available. Although the high-purity magne- 
sias are more expensive than the lower grade magnesias (<93 pet 
MgO), their improved performance often warrants the higher costs. 

Consumption of refractory-grade MgO has declined in the 
United States and other countries. Bureau of Mines statistics (2) 
show 290,271 st of refractory MgO and 99,517 st of caustic cal- 
cined MgO were shipped and used in the United States in 1985. 
World capacity is estimated to be about 10 million tons, with 80 
pet derived from natural MgCOs. Refractories account for greater 
than 70 pet of the MgO consumed in the United States and the world. 
The decline in world steel production and improvements in refrac- 
tory life have resulted in world overcapacity and the closure of 
several plants. 

Periclase is produced by the crystallization of magnesia and 
the growth and agglomeration of magnesia particles during calcin- 
ing and sintering of magnesite, brucite, or chemically precipitated 
hydroxide. These magnesia particles are referred to as grains rather 
than crystals because they are polycrystalline. 

Generally, seawater and brine magnesias are obtained by using 
suitable alkali materials such as limestone or dolomite to precipi- 
tate magnesium hydroxide. However, during precipitation, impu- 
rities are inadvertendy introduced. For example, seawater magnesias 
often contain boron, the concentration of which depends on the 
boron content of the feed stock or bittern. Magnesia produced from 
brine wells and lake sources generally contains less boron but may 
contain traces of chloride and/or sulfate anions. However, during 
calcination to remove water of hydration, many of these impuri- 
ties are volatilized. Material produced from magnesite ore is usually 
dead burned in rotary kilns at temperatures exceeding 1,600° C 
to remove volatile and low-melting-point impurities. Besides impu- 
rities in the MgO grain, refractory-grade MgO users are concerned 
with the crystallite size and density (larger crystallite size and higher 
density giving greater refractory life) (3). 



The Bureau characterized a series of 13 magnesia raw materials 
(natural and synthetic) as to chemistry and mineralogy of the fired 
periclase (4). The results of these tests are summarized in tables 
2 and 3. 

Regardless of the magnesia source, processing variables dur- 
ing precipitation, sizing, pelletization, and calcination often affect 
the physical properties and the integrity of the refractory grains 
produced. Several investigators (5-7) have reported that modulus 
of rupture (MOR) data obtained at elevated temperatures gave the 
most reliable indication of the high-temperature strength and per- 
formance of a refractory. 

Using the periclase materials, shown in tables 2 and 3, sam- 
ples of each magnesia grain were evaluated by the Bureau. Hot MOR 
determinations were made at 1,500°, 1,550°, and 1,600° C on 
U- by U- by 2-in samples. The test results indicated that (1) the 
seawater magnesias generally had the lowest hot strengths while 
a brine periclase material had the highest hot strength, (2) an opti- 
mum CaO-SiOj (C-S) ratio with reference to hot strength existed 
for each periclase material, and (3) B2O3 contents over 0. 1 pet dra- 
matically lowered the hot strength of periclase grains. 

The effect of controlled additions of selected metal oxides, as 
well as the adjustment of the C-S ratio, on the hot strength of 
periclase grains was investigated in a followup study (8). Oxides 
of Zr, Ti, Ta, Sc, Fe, and Mn were added in amounts up to 2 pet 
to three periclase grains. The C-S mole ratios were adjusted to values 
of 1.5, 2.0, 2.5, and 3.0. Test results showed that threefold to four- 
fold increases in hot MOR of some periclase materials were obtained 



Table 2.— Chemical analysis of commercial 
periclase grains, weight percent 



Sample type MgO 



B2O: 



2<-'3 



CaO 



SiO, 



AI2O3 



FezOa 



Brine 93.7-99.2 0.0 -0.09 0.6 -2.8 0.1-1.9 0.04-0.26 0.08-0.25 

Magnesite. 88.2-93.5 .01- .03 2.8 -4.9 1.7-3.2 .23- .34 .36- .82 
Seawater.. 93.3-97.6 .01- .33 .53-1.2 .6-3.3 .26- .38 .27- .7 



Table 3.— Mineralogy and physical properties 
of commercial periclase grains 

Brine Magnesite Seawater 

Mineralogy, X-rayi MjS, CjS, C3MS2, C2S, MjS, CMS 

C3MS2 aAljOa, 

CMS 

Bulk density g/cm^ . . 3.23-3,37 3,11-3.32 3.13-3.41 

Average grain size nm . . 20-60 30-60 30-60 

1 M2S = forsterite, CMS = monticellite, C3MS2 = merwinite, C2S = dical- 
cium silicate. 



30 



with additions of Zr02 and by adjustment of their C-S ratio to 2.5 
to 3.0. Results also showed that for each periclase refractory raw 
material, there existed a combination of percent Zr02 addition and 
C-S ratio adjustment to attain optimum hot MOR values. Data also 
indicated that the improved hot strengths can be attributed to the 
high-temperature secondary phase (dicalcium silicate) formed 
between the periclase impurities and the additives, resulting in 
improved intergranular high-temperature bond strength. 

The effects of chemical adjustments of periclase grain and 
adjustments to an Mg(OH)2 slurry prior to grain densification were 
investigated (9) using full-size commercially processed brick sam- 
ples. Strength values, both hot and cold, for the optimized brick 
mixes were superior to those for a commercial 98-wt-pct-MgO 
refractory. The best results were obtained for a natural magnesite 
with an adjusted C-S ratio of 3.0 and a 1 .0-wt-pct addition of ZrOj, 
and for a seawater periclase with an adjusted C-S ratio of 2.5 and 
a 0.5-wt-pct addition of Mn02. For samples produced from chem- 
ically modified Mg(OH)2 slurries, additions of both Mn02 and Zr02 
along with an increase in C-S ratios produced significant increases 
in hot MOR, similar to those resulting from chemical additions made 
to periclase grains. The most effective modifications were addi- 
tions of Mn02 to brine-derived periclases having a C-S ratio adjusted 
to 3.0. These results on chemically modified magnesia refracto- 
ries indicated that such material could potentially substitute for 
magnesia-chrome refractories containing imported chromite. 

A study (10) was also made to determine if the physical prop- 
erty improvements noted in periclase through metallic oxide addi- 
tions of powders would occur in fired brick samples impregnated 
with water-soluble additions of metallic salts. Samples of 90- and 
98-pct-MgO brick were soaked in solutions containing Al, Ca, Cr, 
Co, Fe, Mg, Mn, Mo, Ni, Si, Sr, Sn, Ti, or Zr ions. Additions 
to 98-pct-MgO brick did not generally result in statistically signifi- 
cant property improvements. Statistically significant improvements 
were noted in MOR, slag resistance, and spalling resistance of the 
90-pct-MgO brick for additions of Al, Mg, or Sn. High-temperature 
performance of treated 90-pct-MgO refractories was equal to or 
approached that of 98-pct-MgO brick. The marked improvements 
due to Sn were obtained with additions of only 1 . 14 wt pet Sn02. 



DOLOMITE 

Dolomite (11) (CaCOj-MgCOa), idenfified by Dolomieu in 
1791, occurs as sedimentary deposits similar in nature to limestone. 
Geologically, some dolomites are precipitated directly from sea- 
water, but most dolomites are a result of the alteration of calcium 
carbonate sediments or rocks by hypersaline brines. Good exam- 
ples are the almost-pure dolomite Silurian reefs in northern Illinois, 
Indiana, and Ohio, and in southern Michigan. Other carbonate 
minerals are found associated with dolomite, but usually not in great 
quantities. 

The Bureau estimated that 31 million st of dolomite was 
produced by 60 companies in 24 States during 1985. The use of 
dolomite in refractories is minor compared with the total amount 
produced. More than three-fourths of the dolomite quarried in the 
United States is used as an aggregate or a soil conditioner. In 1985 
(12), the amount of dead-burned dolomite sold or used by producers 
was 378,000 st, which was only 2.4 pet of the total volume of lime 
sold or used by producers. 

The only States mentioned by Colby (13) as producing 
refractory-grade dolomites were Alabama, California, Colorado, 
Dlinois, Michigan, Nevada, Ohio, Pennsylvania, Utah, and West 
Virginia. Impurities in refractory-grade dolomite are typically less 
than 2 pet (14). Worldwide occurrence of dolomite deposits with 
good commercial value (good chemical and physical properties 



located close to the consumer) for refractories or flux usage are 
rare (15). 

Dolomite used as refractories has the CO2 driven off and the 
grain densified (called dead burning) to increase the calcium sta- 
bility. For high-purity, low-iron dolomite, this requires heating 
to 1 ,800° C. Lower purity dolomite for fettling purposes (the proc- 
ess of repairing furnace bottoms with loose dolomite grain) is gener- 
ally sintered in the 1,400° to 1,600° C range. Dead-burned dolomite 
is generally purer than fetding grades. Iron oxide is sometimes inten- 
tionally added to the fettling grade to meet the special use require- 
ments. As open-hearth steelmaking decreases worldwide, the need 
for fettling grades of dolomite will decline to zero. The open hearth 
was once the most important use area for refractory-grade dolomite. 

The main applications of dolomite brick in the steel industry 
are in BOF furnaces, ladle linings, AOD vessels, and some EAF 
use. Ladle use of dolomite brick is not the largest consumption area 
in steelmaking, but it has the most promising future. Cleaner, higher 
quality, and more cost-effective steel is claimed to be produced by 
dolomite brick in ladles as more secondary steel processing occurs 
there. The output of dead-burned dolomite refractories, however, 
suffers from reduced steel output. 

Dolomite bricks have a stability problem (the CaO reacts with 
water or CO2, causing brick degradation) that can be overcome by 
special handling procedures or by coating the material with tar or 
pitch. Impurities such as silica can also react at high temperatures 
with lime in dolomite to produce beta-dicalcium silicate. On cool- 
ing, this converts via a large volume expansion to gamma-dicalcium 
silicate, which can cause dusting of a brick. 

The overall chemistry as well as the ratio of accessory oxides 
to the combined MgO and CaO content affect the physical and the 
chemical resistance of dead-burned dolomite grains. Since the 
majority of dolomite grains are used in the form of organically 
bonded brick or specialty mixes, this is the logical form in which 
to measure hydration resistance. Hubble (16) devised a hydration 
test that led to the establishment of a standard test, ASTM C492-66 
(17). 

The American Society for Testing and Materials (18) classi- 
fies dolomite refractory raw materials as (1) raw refractory dolo- 
mite, (2) calcined refractory dolomite, and (3) dead-burned 
refractory dolomite. This classification is based primarily on MgO 
content, loss on ignition, and impurity contents. 

Samples of 14 different raw dolomites from sources in Ala- 
bama, Ohio, Pennsylvania, Missouri, Michigan, California, and 
Wisconsin were evaluated (19). Prior to this study there were no 
published data on the refractory properties of domestic dolomite 
even though it was receiving serious consideration as a substitute 
for periclase. Eight of these samples were obtained from suppliers 
of refractory-grade dolomites; the other six were representative of 
dolomites that were used for nonrefractory applications. The raw 
materials were characterized as to chemical, physical, and thermal 
properties. These test results are summarized in table 4. All of the 
materials contained at least 49 wt pet combined MgO and CaO. 
Raw apparent specific gravities ranged from 2.81 to 2.87 and the 
raw bulk densities ranged from 2.55 to 2.80 g/cm^ The major acces- 
sory minerals associated with these dolomites were calcite and 
quartz. 

The thermal analyses of the materials were characterized by 
two endothermic peaks, one occurring between 780° and 820° C 
and the other occurring between 860° and 920° C. Examination 
of thin section photomicrographs of the raw dolomites indicated 
that the average crystallite grain size ranged from around 100 nm 
up to about 750 ^m. The microstructures of two Pennsylvania dolo- 
mites that are suitable for calcining to high-density dead-burned 
grain in a single firing were characterized by the largest average 
crystallite grain sizes and by a large number of twinned grains. It 
is possible that the large grain sizes and occurrence of twinned grains 



31 



Table 4.— Properties of raw domestic dolomites 



State of source 




Chemical analysis, 
wt pet 






Loss on 
ignition 


Apparent 
specific 




MgO 


CaO 


SiOa 


AI2O3 


FezOa 


gravity 


Alabama 

Ohio 

Pennsylvania 

Michigan 

Missouri 


20.1-20.8 

19.5-21.2 

21 -21.3 

21 -21.2 

19.2 

21.2 

21.7 


30.1-30.5 

29.6-30.6 

27.6-30.8 

30.3-30.6 

31.2 

30.8 

31.1 


1.1 -1.5 
.02-1.7 
.1 - .3 
.5 
.3 
.3 
.5 


0.4-0.8 

.1- .8 

.1- .2 

.1 

.1 

.04 
.1 


0.2-0.3 

.1-3 

.2- .4 

.1- .2 

3.6 

.2 

.1 


46.5-47.3 

45 -47.5 

46.4-47.1 

47.2-47.4 

45.4 

46.9 

45.9 


2.85-2.87 

2.84-2.87 

2.81-2.86 

2.84 

2.84 


Wisconsin 

California 


2.86 
2.82 



has some influence upon the calcination and densification of these 
dolomites. These Pennsylvania dolomites, along with materials from 
Missouri, California, and Ohio, were found to be suitable for refrac- 
tory use. 

FLAKE GRAPHITE 

Continuous casting of steel combined with increased operat- 
ing temperatures and demands for longer refractory life have focused 
attention on carbon-containing refractories (20-21). Carbon, in the 
form of natural flake graphite, imparts a high degree of oxidation 
resistance, reduces the level of wettability by slag, and increases 
the thermal conductivity of the refractory. Graphite-base refracto- 
ries are used in BOF, EAF, and transfer ladles, as well as pouring 
tubes and nozzles. The carbon-base refractories typically contain 
MgO, AI2O3, or dolomite bonded by pitch or resin with up to 30 
wt pet natural flake graphite. The natural flake graphite used in 
refractories is totally imported. Deposits of natural flake graphite 
are available in the United States, but they are not economical 
to mine. 

Flake graphite occurs as flat, platelike particles occurring in 
layers of metamorphosed silica-rich sedimentary rock such as quartz 
and mica schists, felspathic and micaceous quartzites, gneisses, and 
marble. Commercial importance varies considerably, depending 
upon flake size and carbon content (22). Import of natural flake 
graphite for refractory consumption has held fairly constant and 
was about 1,000 tons in 1980, the last year the Bureau of Mines 
reported this figure. 

Graphite, because of its unique properties, makes substitution 
of other materials difficult. Although carbon blocks are used in blast 
furnaces, synthetic graphite is considered too expensive to be sub- 
stituted for graphite use. For many uses, graphite from different 
origins is not considered interchangeable. Many consumers use 
blends of graphites from different sources to eliminate supply 
problems. 

Mag-carbon bricks, developed in the United States, were 
applied to a large number of EAF furnaces in Japan. The current 
trend is for EAF steelmakers to replace chemically bonded magnesia 
brick with more durable resin-bonded mag-carbon brick with car- 
bon contents of 15 to 20 pet. Furnace service life is about 500 heats. 
EAF operating temperatures can be raised to over 2,000° C utiliz- 
ing water-cooled sidewalls. In BOF's, brick linings containing up 
to 20 pet natural flake graphite have a lining life of around 750 
casts. Mag-carbon bricks are also used in ladles, with a carbon con- 
tent of about 10 pet. Lining life, because of the oxidizing 
atmosphere, is approximately 50 heats. 

The optimum ash content for graphites used in mag-carbon 
brick is 2 pet, although graphites with ash contents as high as 10 
pet are used. Research has shown that silica, alumina, and iron impu- 
rities in graphite form low melting compounds that shorten brick 
life. Flake size in mag-carbon refractories ranges between 150 and 
710 /tm. The relationship of flake length to width should be greater 
than 20:1 in order to minimize oxidation. 



RG 



0.1 


12.6 


31.2 


47.9 


7.6 


.5 


.1 


.8 


23.4 


21.9 


.9 


.7 


2.2 


43.75 


NA 



Table 5.— Natural flake graphite properties 

Graphite grade ... 85 90 95 100 

Screen analysis, wt pet 
retained: 

Plus 18 0.2 

Minus 18 plus 30. . . 1.7 14.4 9.8 11.9 

Minus 30 plus 40. . . 23 26.4 25.9 27.7 

Minus 40 plus 60. . 59.7 44.3 48.1 48.1 

Minus 60 plus 80. . . 14.3 12.1 13.8 10.7 

Minus 80 plus 100. . 1.2 1.5 1.6 1.1 

Minus 100 .2 1.3 .9 .3 

Ash wt pet . . 13.7 10.6 6 

Ash chemistry, wt pet: 

SiOz 55.9 45.2 43.8 NA 

FozOa 11.8 23.7 26.1 NA 

AI2O3 26.2 22.8 25.7 NA 

MgO 1.1 6.8 4.3 NA 

CaO 1 1.3 1.2 NA 

B2O3 ND ND ND NA 

Ash, PCE 14-15 6-7 6 NA 



NA Not analyzed. ND Not detected. 

The need to develop a substitute material for natural flake graph- 
ite in basic refractories has been recognized (23). However, fun- 
damental high-temperature engineering data, such as are available 
on mag-carbon systems, does not exist for dolomite-carbon refrac- 
tories. A study at the Bureau's Tuscaloosa (AL) Research Center 
(24), in cooperation with the J. E. Baker Co., on the properties 
of several flake graphites (listed in table 5) was initiated. The grade 
100 is a thermally purified material, and the grade RG is a mate- 
rial treated with boron to increase oxidation resistance. 

The five grades of natural flake graphite were used to evaluate 
the effects of different amounts of graphite on the physical and ther- 
mal properties of dolomite-carbon refractories containing 4 pet phe- 
nolic resin binder. Refractory formulations were mixed as 150-lb 
batches in a high-speed countercurrent mixer. Brick of 6- by 9- 
by 3'/2-in dimensions were pressed at 20,000 psi and cured on a 
commercial schedule. One set of samples contained 10-wt-pct addi- 
tions of the five different graphites, and another set contained 
increasing quantities (0, 5, 10, 15, 20, and 30) of a grade 90 graph- 
ite. Changes in oxidation resistance at 1 ,093 ° C, hot strength from 
260° to 1,510° C, and deformation under load at 1,500° C were 
determined. 

Results indicated that carbon purity of 10-wt-pct-graphite addi- 
tions did not influence hot strength or deformation under load. When 
the quantity of grade 90 natural flake graphite addition varied 
between and 30 wt pet, hot MOR was highest and deformation 
under load lowest with 10-wt-pct additions. As the hot MOR test 
temperature was increased from 260° to 1,510° C, the strength 
difference observed between 0- and 30-wt-pct additions of a grade 
90 graphite became less. Additions of a boron-treated graphite 
caused significantly higher strength and lower deformation under 
load at high temperatures, while an addition of ball clay signifi- 
cantly reduced strength and increased deformation. 



32 



A subsequent study is being conducted to evaluate substitute 
syntiietic carbon materials for natural flake graphite in dolomite- 
carbon refractories. Polyacrylonitride and pitch-carbon fibers of 
Ys- and Vi6-in length and carbon flakes punched out of synthetic 
graphite paper were tested. Commercial -grade refractory mixes 



were prepared and pressed into bricks. Carbon fiber additions above 
2 pet resulted in rebound or springback during pressing. Two- 
percent carbon fiber additions resulted in hot strengths of 260 psi 
versus 600 psi for samples with natural flake graphite additions. 



REFERENCES 



1. Van Dreser, M. L., and J. E. Neely. Refractories for Steelmaking 
in the U.S. A . . . Current Practice and Future Trends. Pres. at First Inter- 
national Conference on Refractories, Tolcyo, Japan, Nov. 15, 1983, 28 pp.; 
available upon request from T. A. Clancy, BuMines, Tuscaloosa, AL. 

2. Kramer, D. A. Magnesium Compounds. Ch. in BuMines Minerals 
Yearbook 1985, v. 1, 1987, pp. 661-667. 

3. Coope, B. Magnesia. Sec. in "IM" Refractories Survey 1986— Raw 
Materials for the Refractories Industry, ed. by E. M. Dickson, 1986, 
pp. 28-42. 

4. McLendon, J. T., N. S. Raymon, and H. Heystek. Relationship of 
Mineralogical and Chemical Composition of Refractory Periclase to Modulus 
of Rupture at 1,500° to 1,600° C. BuMines RI 8386, 1979, 18 pp. 

5. Buist, D. S., A. Hatfield, and H. Pressley. Modulus of Rupture as 
an Index of Potential Refractory Performance in Service. Trans. J. Br. 
Ceram. Soc, v. 68, 1969, pp. 45-47. 

6. Gilpin, W. C, and D. R. F. Spencer. New Developments in Dead- 
Burnt Magnesite and Dead-Burnt Dolomite. Refrac. J., v. 47, No. 4., Apr. 
1972, pp. 4-16. 

7. Jackson, B. , and J. Laming. The Significance of Mechanical Projjcr- 
ties of Basic Refractories at Elevated Temperature. Trans. J. Br. Ceram. 
Soc, v. 68, 1969, pp. 21-28. 

8. Raymon, N. S. Influence of Selected Additives and CaOiSiOj Ratio 
on High Temperature Strength of MgO Refractories. BuMines RI 8732, 
1982, 8 pp. 

9. Bennett, J. P., and T. A. Clancy. Magnesia Refractories Produced 
From Chemically Modified Periclase Grains and Mg(0H)2 Slurries. BuMines 
RI 8848, 1984, 14 pp. 

10. Bennett, J. P. High-Temperature Properties of Magnesia-Refractory 
Brick Treated With Oxide and Salt Solutions. BuMines RI 8980, 1985, 1 1 pp. 

11. Carr, D. D., and L. F. Rooney. Limestone and Dolomite. Ch. in 
Industrial Minerals and Rocks. AIME, 1975, 1360 pp. 

12. Pelham, L. Lime. Ch. in BuMines Minerals Yearbook 1985, v. 1, 
1987, pp. 635-644. 



13. Colby, S. F. Occurrence and Uses of Dolomite in the United States. 
BuMines IC 7192, 1941, 21 pp. 

14. Hopkins, D. A. Dolomite. Am. Ceram. Soc. Bull., v. 66, No. 5, 
1987, pp. 759-760. 

15. Dickson, T. Dolomite. Sec. in "IM" Refractories Survey 1986— Raw 
Materials for the Refractories Industry, ed. by E. M. Dickson, 1986 
pp. 44-46. 

16. Hubble, D. H. , and W.J. Lackey. Hydration Test for Dead-Burned 
Dolomite. Am. Ceram. Soc. Bull., v. 41, No. 7, 1962, pp. 442-446. 

17. American Society for Testing and Materials. Standard Test Method 
for Hydration of Granular Dead-Burned Refractory Dolomite. C492-66 in 
1981 Annual Book of ASTM Standards: Part 17, Refractories, Glass, 
Ceramic Materials; Carbon and Graphite Products. Philadelphia, PA, 1981, 
pp. 404-405. 

18. Standard Classification of Granular Refractory Dolo- 
mite. C468-70 in 1981 Annual Book of ASTM Standards: Part 17. Refrac- 
tories. Glass Ceramic Materials; Carbon and Graphite Products. Philadel- 
phia, PA. 1981, pp 383-384. 

19. Clancy, T. A. Dolomite Refractories, and Their Potential as Substi- 
tutes for Imported Chromite. BuMines IC 8913, 1983, 18 pp. 

20. Brown, A. The Properties of Ceramic Graphite Bodies. Refrac. J., 
No. 2, Mar./Apr. 1985, pp. 7-10. 

21. Cooper, C. F. Refractory Applications of Carbon. Trans. J. Br. 
Ceram. Soc, v. 84, 1985, pp. 48-53. 

22. Kenan, W. M. Graphite. Am. Ceram. Soc. Bull., v. 66. No. 5, 1987, 
pp. 762-763. 

23. Hayashi, T. Recent Trends of Refractory Technologies in Japan. Tai- 
kabutsu Overseas, v. 4, No. 1, 1984, pp. 3-19. 

24. Bennett. J. P. The Effect of Different Natural Flake Graphite Addi- 
tions on the High-Temperature Properties of a Dolomite-Carbon Refrac- 
tory. BuMines RI 9111, 1987, 12 pp. 



33 



BASIC RESEARCH ON CORROSION OF IRON-BASED MATERIALS 



By David R. Flinn^ 



ABSTRACT 



The Bureau of Mines has conducted corrosion-related research at a significant level for the 
past three decades. Much of this research consisted of exposure testing to evaluate commercially 
available metals and alloys for use in minerals and metals processing environments and to deter- 
mine the corrosion behavior of newly available metals such as Ti, Nb, Ta, and Zr in selected environ- 
ments. In recent years a significant portion of the research has been redirected toward the 
development of a fiindamental understanding of the roles of alloy composition and microstructure 
and of the process environment in determining corrosion behavior. This paper discusses the impor- 
tance of this understanding to the efficient utilization of mineral resources and examples are given 
of research designed to accomplish these objectives. 



INTRODUCTION 



Corrosion is the degradation of the mechanical, physical, 
and chemical properties of materials resulting from chemical- 
electrochemical processes and the interaction of these processes with 
mechanical and wear processes. The cost of corrosion to the U.S. 
economy was estimated to be $70 billion in 1975, or about 4.2 pet 
of the gross national product (1).^ 

With current technology, approximately 15 pet of that cost could 
have been avoided and the opportunity for even greater savings exists 
with the development of new and improved corrosion control and 
prevention measures. The benefits of corrosion-related research and 
development activities are well recognized in terms of the savings 
in energy and materials to reduce excess capacity, enhance produc- 
tion, reduce product losses, lower maintenance, and improve safety. 
In a recent report by the National Materials Advisory Board (2), 
the value of corrosion control was estimated "... to have an eco- 
nomic impact from fifty- to a hundred-fold greater than its dollar 
value." A further benefit to be expected from the development of 
corrosion control technologies is the conservation of critical and 
strategic materials by substitution of more abundant and domesti- 
cally available materials and by improving service life in environ- 
ments used in existing technologies and the more severe 
environments of new and emerging technologies. 

In a recent report (3) published by Battelle Columbus Labora- 
tories and based on a conference of materials engineers, corrosion 



' Supervisory research chemist, Albany Research Center, Bureau of Mines, 
Albany, OR. 

^ Italic numbers in parentheses refer to items in the list of references at the end 
of this paper. 



scientists, and electrochemists on corrosion control and prevention, 
60 pet of the highest priority research needs applied to corrosion 
problems experienced in the minerals and materials processing 
industries. The same report (3) pointed out that, for the primary 
metals industry, corrosion problems are usually sidestepped by oper- 
ating processes at reduced energy efficiency. 

An important aspect of the Bureau of Mines mission is the 
development and assessment of technological alternatives in mineral 
processing and related materials development to support the secu- 
rity and economic well-being of the Nation with regard to mineral 
and material processing, performance, utilization, and recycling. 
The Bureau conducts long-range and generic basic research on cor- 
rosion phenomena important to all minerals and materials process- 
ing industries. The objective of this research is to provide a scientific 
understanding of corrosion processes, including the roles of mate- 
rial composition and structure as well as the role of the environ- 
ments to which the material is likely to be exposed. Through this 
understanding it is possible to determine cost-effective corrosion 
control strategies that may involve such options as materials sub- 
stitution or materials protection. This research has the goal of provid- 
ing the underpinning for the development of improved and new 
technology for the minerals and materials processing industries. 
A further goal is the promotion of conservation of critical and stra- 
tegic materials throughout the industrial community. These goals 
are reflected in the many accomplishments of the Bureau in corro- 
sion research achieved over the past three decades (4). Most of this 
research has been concerned with aqueous environments. The work 
discussed here will be limited to aqueous environments and to iron- 
based alloys. 



34 



ROLE OF ALLOY COMPOSITION AND STRUCTURE 



Except for comparatively rare instances, most metals and alloys 
are not thermodynamically stable with respect to even the most com- 
mon environments. Fortunately, some of these metals and alloys 
can be used in many environments because they corrode at low rates. 
The chemical composition of an alloy strongly influences its cor- 
rosion resistance. Tomashov (5) has reviewed the mechanisms by 
which different alloying elements can affect corrosion resistance. 
Some elements increase thermodynamic stability, some retard 
cathodic reactions (hydrogen evolution, oxygen reduction) that are 
required to support the anodic alloy dissolution, while others retard 
the anodic dissolution reaction itself (5). 

This latter mechanism, which usually involves the formation 
of a protective ("passivating") surface film, is the most common. 



and scientifically most interesting, way that alloys resist dissolu- 
tion. Reviews of the processes thought to be involved in the for- 
mation of protective passive films on alloys are widely available 
in the literature (as examples, see references 6 through 8). Even 
with this knowledge, there remains a very incomplete understand- 
ing of how the elemental composition and the structure 
(homogeneity, degree of crystallinity, defect structures, etc.) of an 
alloy influence the formation of protective passive films. An impor- 
tant research area of continuing interest in the Bureau is to gain 
a fundamental understanding of how the composition and structure 
of an alloy influence the formation of corrosion films that can pro- 
vide resistance to aggressive environments. 



BUREAU OF MINES FUNDAMENTAL CORROSION RESEARCH3 



EFFECTS OF ENVIRONMENT 

Corrosion is an electrochemical process involving sf)ecies from 
the alloy and from the environment. For this reason, meaningful 
research on corrosion processes can only be accomplished when 
the chemistry of both the alloy and the environment are controlled 
and/or characterized. The importance of a controlled environment 
was demonstrated for electrochemical polarization studies of iron 
in sulfuric acid (9). The electrochemical corrosion behavior of iron 
in solutions made from distilled water and reagent-grade sulfuric 
acid was shown to be irreproducible. For some tests, a potential 
range where the iron was passive would be observed, while no pas- 
sivation was observed in other, apparently identical, tests. When 
special precautions were taken to use high-purity deionized water 
and special high-purity acid, the polarization studies yielded very 
consistent results. In the same study (9), solution purity was found 
to have much less influence on the electrochemical corrosion 
behavior of Fe-18Cr. 

In a later study (10), the electrochemical behavior of Fe-18Cr 
was investigated in an even higher purity environment. Very care- 
fully purified solutions and gases, gaslight cells, and a special high- 
purity Fe-18Cr alloy wire were used. In this ultra-high-purity sys- 
tem, the Fe-18Cr alloy did not corrode in a lA^ H2SO4, H2 satu- 
rated solution, but instead attained a potential very near that for 
a reversible hydrogen electrode. This potential was maintained for 
months in the solution. 

It was determined (10) that the alloy could be caused to cor- 
rode by cathodically polarizing the electrode; the Fe-18Cr alloy 
could be returned to the noncorroding state by a brief anodic polar- 
ization of the electrode. Specially developed electrochemical tech- 
niques were used in this study (10) to prove that metal-oxygen 
species were not present in significant amounts on the electrode 
surface and could not be the cause of the behavior observed. 

These two studies (9-10) are examples of why it is important 
to characterize and/or control the materials and the environment 
being used. Although such extreme purification procedures do not 
represent realistic situations, they do provide important baseline 
information that can be used to evaluate the effects of controlled 



' As will be demonstrated in the examples that follow, the understanding of corro- 
sion processes cannot be separated from the environments and materials involved. 
For this reason, the fundamental corrosion research effort in the Bureau is multidis- 
ciplinary, involving chemists, physicists, engineers, metallurgists, and materials 
scientists. 



additions of species, such as oxygen gas or chloride ion, that could 
be expected to be present in many important aqueous environments. 

A more recent study was conducted to determine the effect of 
oxygen on the formation of passive films on stainless steel at open 
circuit conditions (11). An Fe-18Cr alloy was selected for the study 
because it does not spontaneously passivate from a state of corro- 
sion when oxygen is added to the environment, but it can be passi- 
vated by electrochemical polarization. Because of these 
characteristics it was possible to determine the amount of dissolved 
oxygen needed in lA^ H2SO4 to maintain passivity following con- 
trolled potential passivation. In addition, the composition of the pas- 
sive film as a function of potential was determined using surface 
analytical techniques. The effect of solution oxygen partial pres- 
sure on the potential-time relationship of Fe- 1 8Cr was examined 
by first passivating the alloy for 80 min at 0.6 V versus the normal 
hydrogen electrode (NHE) potential in nitrogen-saturated (IN) 
H2SO4, and then simultaneously releasing the potential to open cir- 
cuit and substituting various O2-N2 gas mixtures for the nitrogen. 

The results of these tests are shown in figure 1 . For solution 
oxygen partial pressures below 0.045 atm (1 .7-ppm O2 solution con- 
centration) the potential of the passivated alloy was observed to fall 
into the range of active corrosion after about 2x10' min. For solu- 
tion oxygen partial pressures above this value, the open circuit poten- 
tial of the alloy was observed to decrease to a minimum and then 
increase again. At times in excess of 10* min (not shown), the poten- 
tial leveled off near 0.6 V. The 1. 7-ppm oxygen concentration 
appears to be the minimum amount required to maintain the alloy 
in the passive state under the experimental conditions used (11). 
The mechanism for breakdown of these passive films probably 
involves trace impurities, so that the exact level of oxygen required 
to maintain the passive film wUl differ for various solutions purities. 

Also discussed in this paper (11) were the effect of time and 
potential on the composition of the passive film as determined by 
Auger electron spectroscopy (AES). Fe-18Cr was electrochemi- 
cally passivated at 0.6 V for 80 min in oxygen-saturated solution 
(approximately 29-ppm O2 concentration) and released to open cir- 
cuit. The chromium atomic fraction compared to iron in the corro- 
sion film was found to remain at approximately 0.45 over the 
potential-time region shown in figure 1 up to approximately 10' 
min. At longer times and higher potentials, the chromium fraction 
in the fdm increased rapidly, approaching 0.65 at longer times. 
These fmdings were also supported by X-ray photoelectron spec- 
troscopy (XPS) analyses (11) that showed a decrease in the iron 
content of the film with increasing potential and a simultaneous 



35 



decrease in film thickness. These results demonstrate the dynamic 
nature of the passive film on chromium-containing iron-based alloys. 
They indicate that the behavior of many stainless steels in acidic 
environments may well be as dependent on such variables as solu- 
tion oxygen content and impurity levels as on any chemical or elec- 
trochemical treatment of the alloy. 

The effect of chloride ion on the passive state of Fe-18Cr and 
of AISI Type 430 stainless steel (430SS) was determined in a study 
(12) similar to that described previously. The alloys were elec- 
trochemically passivated at 0.6 V NHE for 80 min in oxygen satu- 
rated lA^ H2SO4. After the period of passivation the alloy was 
released to open circuit, and a behavior similar to that shown in 
figure 1 was observed. At selected times chloride ion (as NaCl), 
in amounts ranging from 5 to 44 ppt CI", were added and the 
potential -time response was observed. The time difference between 
adding the chloride ion and reaching a state of corrosion is defined 
as the induction time. These times are shown schematically in fig- 
ure 2, where tg and tc are the induction times for the addition of 




10 



Figure 1.— Effect of oxygen partial pressure on potential vs 
time behavior of Fe-18Cr samples released to open circuit after 
being passivated for 80 min at 0.6 V in IN H2SO4 at 30° C. Data 
from Covino (11). 



0.6 



~i \ 

CI"addition to 
region C 




Potential 
> behavior 

\ following 
, Cr addition 



"PP \ 

Corrosion potentiol 1 



JL 



10"' 10° 10' 10' 10' 

TIME, min 



10* 



10' 



Figure 2.— Schematic showing time of CI addition and its effect 
on potential-time behavior of passivated Fe-18Cr and 430SS. 
Epp is the primary passivation potential. Data from Covino (12). 



chloride ion in the minimum potential region and in the increasing 
potential region, respectively. It was found, as expected, that 
increasing amounts of chloride ion decreased the induction times. 
However, very significant increases in induction time were observed 
in cases where the chloride ion was added after approximately 100 
min. In the previously mentioned study (11) it was found that pas- 
sive film became thinner and richer in chromium at times in solu- 
tion of 100 min or longer. Investigations of the type reported here 
(11-12) provide significant insight into the roles that environmen- 
tal species play in the formation and breakdown of protective pas- 
sive films. 



EFFECTS OF ALLOY COMPOSITION AND 
STRUCTURE 

The Bureau has conducted detailed studies of the effects of 
alloying elements on the corrosion behavior of steels. Ion implan- 
tation was one method used extensively for preparing iron-based 
alloys containing preselected amounts of the desired alloying ele- 
ment (13-18). 

Ion implantation is a process whereby ions of the desired alloy- 
ing element are accelerated in a vacuum to energies in the range 
of 20 to 100 keV and allowed to impact a target of the base metal 
(13). The ions penetrate this target to a mean depth on the order 
of a few hundred angstroms, with an approximately Gaussian depth 
distribution. This technique offers numerous advantages over other 
alloy preparation methods, including speed of preparation and rela- 
tively low cost for small amounts of experimental alloys, the very 
small amount of alloying element required, the wide range of alloy 
concentrations that can be formed, the preparation of alloys with 
optimum surface and bulk properties for a given application, and 
the preparation of certain metastable alloys that cannot be formed 
by conventional techniques. These attributes make ion implanta- 
tion a very useful tool for preparation of model alloys for corro- 
sion research. 

In the 1970's the Bureau conducted pioneering studies on the 
use of ion implantation as a technique for the preparation of alloys 
for corrosion studies. It was proven (14-15) that nickel or chro- 
mium implanted into iron produced Fe-Ni or Fe-Cr alloys having 
corrosion properties identical to those of bulk alloys of similar com- 
position. Because of the thinness of the ion implanted layer, these 
"surface" alloys obviously would lose their corrosion resistance 
if exposed to aggressive environments for extended times. Even 
so, for the period of time that this alloy layer is intact, it can be 
expected to behave in a manner similar to that of a bulk alloy of 
the same composition. 

In another study (16), it was shown that chromium-implanted 
iron exposed to air at 320 ° C oxidized at a rate identical to an equiva- 
lent bulk alloy for times as great as 1,000 h, and that the oxide 
thickness was identical on the two alloys. In the same study (16) 
it was shown that aluminum implantation into titanium produced 
an alloy that, when coupled to aluminum in a NaCl solution, greatiy 
reduced the galvanic corrosion loss of aluminum compared to that 
produced when aluminum was coupled to unimplanted titanium. 
The implanted aluminum was found to reduce the rate of oxygen 
reduction on the titanium by as much as a factor of 40. 

The effects of ion implantation on stress corrosion cracking 
(17) and on corrosion fatigue (18) of iron-based alloys have also 
been examined. Very significant effects were noted on the stress 
corrosion cracking (SCC) behavior of AISI Type 316 stainless steel 
(316SS) in boiling MgClz by implanted Si, N, or Ar. Argon implan- 
tation reducted the time to failure by 30 to 40 pet compared to unim- 
planted samples for a total implant dose of 3x10'* ions/cm 
(approximately equivalent to 20 at. pet in the implanted volume). 
This detrimental effect of argon was caused by ion damage to the 



36 



surface region during implantation (approximately 40 nm in depth). 
Implantation of nitrogen into the 316SS resulted in a similar decrease 
in time to failure by SCC. 

Examination of the nitrogen-implanted surface by scaiming elec- 
tron microscopy (SEM) following these tests showed that portions 
of the surface had been ruptured along slip planes by an explosive 
escape of gas from below the surface. An elemental depth profile, 
obtained by AES, of the surface of a nitrogen-implanted sample 
after SCC failure is shown in figure 3. A significant amount of nitro- 
gen remained in the sample following the test, indicating that much 
of the nitrogen must have been present within the alloy rather than 
just in grain boundaries. 

Implanted silicon had a much different effect on the SCC 
behavior of 316SS. As shown in figure 4, the resistance to SCC 
increased approximately linearly with the fluence (atoms/cm^) of 
implanted silicon. Examination of the silicon-implanted samples fol- 
lowing exposure to the MgClj showed that the surface was cov- 
ered by a film that was rich in silicon and magnesium. This film 
passivated the surface, reducing the general corrosion rate. Sili- 
con also reduced the SCC crack propagation rate. This effect was 
not anticipated because of the very thin implanted region. 

In another study (18), the effects of implanting Ti or a combi- 
nation of Mo and Ta on the corrosion fatigue behavior of 1018 car- 
bon steel were examined. The Mo plus Ta implanted steels behaved 
no differently from unimplanted steel in the 0. lA^ H2SO4 solution, 
with apparently identical corrosion potentials and surviving about 



UJ 

o 



UJ 

o 

Z) 

< 




50 100 

SPUTTERING TIME, min 



150 



the same number of cycles in a rotating beam test prior to failure. 
The steels were adversely affected by the titanium implantation, 
however. The corrosion rate of the steel was approximately dou- 
bled, and the fatigue life decreased linearly with the titanium implant 
dose. An approximate 15 at. pet concentration of titanium decreased 
the cycles to failure by about 50 pet compared to the unimplanted 
steel . Examination by SEM of the surface of the titanium-implanted 
steel following failure revealed the presence of a film containing 
titanium and iron, with small iron crystallites penetrating this film. 
The reduced fatigue resistance is thought to be caused by an acceler- 
ated corrosion at these small crystallites. 

Another technique that has been used in Bureau research to 
prepare experimental alloys with controlled composition and struc- 
ture is vapor deposition (19). This technique involves sputtering 
a target (or targets) composed of the elements to be incorporated 
in the desired alloy. Energetic inert gas ions are used to remove 
atoms from the target by impact. The sputtered atoms are quenched 
onto a substrate to form the alloy of interest. This quenching proc- 
ess is equivalent to cooling a molten alloy so rapidly that virtually 
no movement of the deposited atoms occurs following deposition 
when the substrate is cold. Depending on such factors as the sub- 
strate temperature and the rate of deposition, the alloy produced 
may be either crystalline or amorphous. 

One alloy system produced by vapor deposition that showed 
extremely good corrosion resistance for a wide range of composi- 
tions was Fe-Zr (19). Vapor deposited Fe-Zr alloys were deter- 
mined to be amorphous (no long-range crystalline ordering) over 
the range of composition from Fe9oZr,o through Fq^^Zt^t. Elec- 
trochemical polarization studies showed the Fe9oZr,o to be approx- 
imately as corrosion resistant as Fe-18Cr in lA^ H2SO4. 




Figure 3. — Elemental depth profile as obtained by Auger elec- 
tron spectroscopy for nitrogen-implanted 316SS indicating 
retained nitrogen after exposure to boiling MgCl2. Data from 
Walters (17). 



1 2 3 

SILICON IMPLANT FLUENCE, 10" atoms/cm^ 

Figure 4.— Relationship of time to failure of 316SS in boiling 
MgClz as a function of amount of silicon implanted. Data from 
Walters (17). 



37 



The electrochemical polarization response for the ¥t-i-iZr(,^ alloy 
is compared in figure 5 to that of Fe, Fe-18Cr, and Zr in 1^112804 
(19). Over the range of potentials from open circuit corrosion (ca. 
-0.2 V) to the high positive potentials where oxygen evolution occurs 
(ca. 1.2 V), the FejaZrev alloy has approximately the same corro- 
sion resistance as zirconium. AES analysis of the Fe-Zr alloys sub- 
sequent to the polarization studies showed that the surface corrosion 
film is composed predominately of zirconium and oxygen, with the 
iron to zirconium ratio in the film being significantiy reduced com- 
pared to its value in the interior of the alloy. 

An additional technique that has been applied to the prepara- 
tion of experimental alloys is laser processing, in which a laser is 
used to melt a thin layer of a surface. When a thin coating is pres- 
ent on the material being processed, the laser beam can provide 
a means of melting and mixing the coating into the surface of the 
substrate. The electrochemical behavior of a series of Fe-Cr alloys 
prepared using a Nd-YAG pulsed laser has recently been reported 
(20). The electrochemical behavior of these alloys exhibited small, 
but important differences from those observed for similar bulk 
alloys. Some, but not all, of these differences were found to be 
caused by the presence of a very protective oxide film on the sur- 
face resulting from the incomplete exclusion of oxygen during the 
laser processing. 

A three-dimensional chromium concentration plot, obtained by 
electron microprobe measurements of a cross section of the laser 
processed region, is shown in figure 6. The original bulk material 
was a Fe-5Cr alloy. Laser mixing of a thin coating with the sub- 
strate produced an approximately Fe-15Cr alloy to a depth of from 
25 to 30 fim. On top of this alloyed region is a chromium-rich oxide 
film, approximately 5 fim thick that was produced during process- 
ing. This oxide layer is clearly visible by optical microscopy or 
SEM examination. Based on the microprobe resolution on the order 
of 1/im, only slight variations in chromium concentration are observ- 
able within the laser-alloyed region. However, chemical etching 
of the cross-sectioned sample resulted in the upper portion of the 
laser alloy, equivalent to the region in figure 6 from approximately 
5 to 15 /im on the depth axis, becoming extremely roughened. From 
approximately 15-^m depth within the alloy and continuing into 
the unalloyed substrate, the etched surface was very smooth, with 
only minor grain boundary dissolution. Because of the large amount 
of irregular dissolution that occurred within grains in the upper por- 
tion of the laser-alloyed region, with features having dimensions 





. 






1.4 


- 


/ ) 






- 




ro 


- 


/^ 


1 


.6 




i'^ 


/ 
Fe/ 


.2 


■^ 1 


/ \ 


■ / 







\ FelSCr 
\ 


/ 
/ 




/ Fe33Zr67/ 


J 


,^^ 


-.2 


- --^- 






-.6 


I 1 1 1 1 1 


1 1 1 1 1 1 


1 1 1 



LOG CURRENT DENSITY, M.A/cm" 



Figure 5.— Polarization curves in deaerated 1N H2SO4 for 
FessZre? amorphous alloy compared to Fe, Fe-18Cr, and Zr. Data 
obtained at polarization rate of 60 V/h. Data from McCormick f79J. 



of the order of 0. 1 /^m, it was determined that an elemental analy- 
sis method having better resolution than the electron microprobe 
(fig. 6) was required to establish whether the laser-alloyed regions 
were compositionally homogeneous. 

Scanning transmission electron microscopy combined with 
energy dispersive X-ray spectroscopy (STEM-EDS) was utilized 
to investigate at high resolution the microstructure and microchemis- 
try within the laser-alloyed region. For this smdy a second alloy, 
also prepared by the laser processing technique, was used. This 
alloy contained approximately 18 wt pet Cr in the laser-processed 
region and the substrate contained 9 wt pet Cr. The alloy and sub- 
strate regions were analyzed as a function of depth into the sample 
for a thinned cross section. For these STEM-EDS analyses, a 
2(K)-nm analysis interval was used with a 50-mn electron beam 
(probe) size. The results of these analyses are shown in figure 7. 




■^rr, 



Figure 6. — Three-dimensional chromium concentration plot of 
a typical laser-processed alloy showing surface oxide, laser- 
alloyed region, and bulk substrate. Chromium concentrations in 
bulk and laser-processed region are approximately 5 and 15 wt 
pet, respectively. Data from Molock (2Q). 



UJ 

o 



o 




1.250 



2,500 



3,750 



5,000 



6,250 



1 1 1 1 1 


500 




.000 


1,500 


2,000 


2,500 


3,0 


— .!_ 




1 




1 
,4 


4 


c 




*" 


1 




1 




* 



500 1,000 

- DISTANCE, nm — 



1,500 



Figure 7.— STEM-EDS chromium concentration profiles. A, Pro- 
file typical of upper portion of laser-processed region, indicat- 
ing extensive microsegregation; B, profile typical of lower portion 
of laser-processed region; C, profile of Fe-9Cr substrate. Data 
from Molock (20). 



38 



By using this very small probe size it was possible to detect 
inhomogeneities in elemental concentration of a few percent over 
a distance as short as 100 nm. The chromium concentration in the 
upper portion of the laser processed region is shown in figure 7 A. 
Referring to figure 6, this analysis trace was taken in a region cor- 
responding on the depth scale from about 7 to 13 /zm. It is evident 
that the chromium concentration is highly erratic in this region, 
and is the cause of the unusual electrochemical polarization and 
chemical etching behavior mentioned earlier. It should be pointed 
out that if the data shown in figure 7 A were grouped in sets of 10 
points and the average concentration was plotted, the result would 
be an apparently uniform chromium level of about 1 8 wt pet for 
the six calculated averages, similar to what was measured by the 



microprobe method. For the lower portion of the laser processed 
region (fig. IB) and the substrate (fig. 7C), the chromium concen- 
tration appears to be uniform based on the STEM-EDS analysis 
using the 50-nm probe diameter and 200-nm sampling interval. 
The cause of the microsegregation of chromium shown in fig- 
ure lA was attributed (20) to impurities migrating in the liquid ahead 
of the solidification front subsequent to the laser processing, and 
to supercooling and formation of nonequilibrium phases in the last 
of the liquid that solidifies. Whatever the cause of this microsegre- 
gation, its effects on the corrosion properties were much more sig- 
nificant than would have been expected considering the scale of 
variation. 



CONCLUSIONS 



During the past few years, several new techniques have become 
available for preparing new alloys with interesting and potentially 
useful properties. There also is an increasing availability of ana- 
lytical instruments having much greater sensitivity and spatial reso- 
lution capabilities. This combination has already resulted in an 
improved understanding of the roles of composition and structure 
in controlling alloy properties and will be of increasing importance 



in the future. These techniques are just as applicable to the improve- 
ment in properties of existing classes of alloys as they are for 
development of new materials. 

Examples of how the improved understanding of corrosion 
processes and materials performance have been applied to real prob- 
lems are given in reference 4. 



REFERENCES 



1. Bennett, L. H., J. Kruger, R. L. Parker, E. Passaglia, C. Reimann, 

A. W. Ruff, H. Yakowitz, and E. B. Berman. Economic Effects of Metal- 
lic Corrosion in the United States, Part I. NBS Spec. Publ. 511-1, 1978, 
65 pp. 

2. National Materials Advisory Board — Commission on Engineering and 
Technical Systems— National Research Council. New Horizons in Elec- 
trochemical Science and Technology (U.S. Dep. of Energy contract B- 
M44SS-A-Z). NAS, Publ. NMAB 438-1, 1986, 147 pp. 

3. Battelle Columbus Laboratories. Research Needs for Corrosion Con- 
trol and Prevention in Energy Conservation Systems (U. S. Dep. Energy 
contract DE-ACO6-76RL01830/MPO-B-F6814-A-K). Feb. 1985, 74 pp. 

4. Flinn, D. R. Optimization of Materials Selection Through Corro- 
sion Science. Paper in Chromium-Chromite; Bureau of Mines Assessment 
and Research. Proceedings of Bureau of Mines Briefing Held at Oregon 
State University, Corvallis, OR, June 4-5, 1985, comp. by C. B. Daellen- 
bach. BuMines IC 9087, 1986, pp. 115-124. 

5. Tomashov, N. D. Corrosion-Resistant Alloys and Prospects for Their 
Development. Protection of Metals (Engl. Transl. of Z. Metallov), v. 17, 
No. 1, 1981, pp. 11-25. 

6. Uhlig, H. H. Passivity in Metals and Alloys. Corros. Sci., v. 19, 
No. 11, 1979, pp. 777-791. 

7. Sato, N., and G. Okanoto. Electrochemical Passivation of Metals. 
Ch. in Comprehensive Treatise of Electrochemistry, ed. by J. CM. Bockris, 

B. E. Conway, E. Yeager, and R. E. White. Plenum (New York), v. 4, 
1981, pp. 193-245. 

8. Hashimoto, K. Passivation of Amorphous Metals. Paper in Proceed- 
ings of Fifth International Symposium on Passivity: Passivity of Metals and 
Semiconductors, ed. by H. Froment (Proc. Conf. Bombannes, France, May 
30-June 3, 1983). Elsevier, 1983, pp. 235-246. 

9. DriscoU, T. J., B. S. Covino, Jr., and M. Rosen. Electrochemical 
Corrosion and Film Analysis Smdies of Fe and Fe-18Cr in IN H2SO4. 
BuMines RI 8378, 1979, 33 pp. 

10. Rosen, M. Electrochemical Corrosion of Iron-Chromium Alloys 
Under Ultra-High-Purity Conditions. BuMines RI 8425, 1980, 66 pp. 



11. Covino, B. S., M. Rosen, T. J. Driscoll, T. C. Murphy, and C. 
R. Molock. TheEffectof Oxygen on the Open-Circuit Passivity of Fe-18Cr. 
Corros. Sci., v. 26, No. 2, 1986, pp. 95-107. 

12. Covino, B. S., and M. Rosen, Induction Time Studies of Fe-18Cr 
and 430SS Under Open Circuit Conditions in Chloride-Containing Sulfuric 
Acid. Corros., v. 40, No. 4, 1984, pp. 141-146. 

13. Sartwell, B. D., A. B. Campbell IH, B. S. Covino, Jr., and P. B. 
Needham, Jr. Characterization of Alloys Formed by Ion Implantation. 
BuMines RI 8434, 1980, 29 pp. 

14. Covino, B. S., B. D. Sartwell, and P. B. Needham. Anodic Polari- 
zation Behavior of Fe-Cr Surface Alloys Formed by Ion Implantation. J. 
Electrochem. Soc, v. 125, No. 3, 1978, pp. 366-369. 

15. Covino, B. S., P. B. Needham, and G. R. Conner. Anodic Polari- 
zation Behavior of Fe-Ni Alloys Fabricated by Ion Implantation. J. Elec- 
trochem. Soc, V. 125, No. 3, 1978, pp. 370-372. 

16. Sartwell, B. D., A. B. Campbell, B. S. Covino, and T. J. Driscoll. 
Applications of Ion Implantation to Metallic Corrosion. IEEE Trans. Nucl. 
Sci., V. NS-26, No. 1, 1979, pp. 1670-1676. 

17. Walters, R. P., N. S. Wheeler, and B. D. Sartwell. The Effects of 
Surface Modification on the Stress Corrosion Cracking Behavior of 316 
Stainless Steel. Corros., v. 38, No. 8, 1982, pp. 437-445. 

18. Sartwell, B. D., R. P. Walters, N. S. Wheeler, and C. R. Brown. 
The Effect of Ion-Implanted Alloy Additions on the Linear Polarization and 
Corrosion Fatigue Behavior of Steel. Paper in Corrosion of Metals Processed 
by Directed Energy Beams, ed. by C. R. Clayton and C. M. Preece. Trans. 
Metall. Soc.-AIME, Warrendale, PA, 1982, pp. 53-73. 

19. McCormick, L. D., N. S. Wheeler, C. R. Molock, and C. L. Chien. 
Corrosion Properties of Amorphous Iron-Zirconium Films in IN Sulfuric 
Acid. J. Electrochem. Soc, v. 131, No. 3, 1984, pp. 530-534. 

20. Molock, C. R., R. P. Walters, and P. M. Fabis. Effect of Laser 
Processing on the Electrochemical Behavior of Fe-Cr Alloys. J. Electrochem. 
Soc, V. 134, No. 2, 1987, pp. 289-294. 



39 



FUNDAMENTALS OF STAINLESS STEEL ACID PICKLING PROCESSES 



By Bernard S. Covino, Jr.^ 



ABSTRACT 



Research on the pickling of stainless steels has been conducted by the Bureau of Mines in cooper- 
ation with the American Iron and Steel Institute (AISI). The objectives of the research were to 
reduce the loss of the strategic and critical metals chromium and nickel, to reduce the use of HNO3 
and HF acids, and to reduce the quantity of waste pickling solutions generated. The model used 
to design the research consisted of removal of scale (pickling) from the stainless steel by under- 
cutting (dissolving) the metal beneath the scale. The dissolution behavior of AISI Type 304 stain- 
less steel (304SS) and of three experimental alloys (Fe-4Cr-13Ni, Fe-12Cr-12Ni, Fe-12Cr-17Ni) 
representative of the chromium-depleted metal beneath the scale was studied as a function of tem- 
perature and the concentrations of HNO3, HF, and dissolved Fe, Cr, and Ni. Results indicated 
that the dissolution process was activation controlled, linearly dependent on HF, Fe, and Cr con- 
centrations, and nonlinearly dependent on HNO3 concentration. A comparison of the behavior of 
304SS to that measured for the experimental alloys indicated that it was possible to optimize the 
pickling of 304SS in terms of metal lost, acids used, and waste generated. A pickling liquor com- 
position of 0.8M-1.3M HNO3 plus 0.5M HF at 50° C can optimize the pickling of 304SS. 



INTRODUCTION 



Some of the basic steps used in the formation of stainless steel 
sheet are given schematically in figure 1 . The hot- and cold-forming 
operations used in this process leave the steel in an unusable work- 
hardened state. A short-term high-temperature annealing operation 
is used to soften the metal to allow ftirther rolling of the metal. 
The effects of annealing are that the bulk microstructure is altered, 
the surface is oxidized to form a thick scale, and the region just 
below the oxidized surface is compositionally altered thus forming 
the chromium-depleted region. While the first effect is desirable 
and controllable, the final two effects are not and result in the loss 
of metal. The oxide scale and chromium-depleted region formed 
during annealing must be totally removed during the pickling oper- 
ation and the metal removed is usually not recovered. In fact, the 
metals lost end up concentrating in the pickle liquor, making it dif- 
ficult to use for ftirther pickling operations, and necessitating the 
disposal of the spent pickle liquor. An in-depth understanding of 
the pickling process and of all of the factors affecting it could help 
to minimize the loss of metals from the stainless steel, the use of 
pickling acids, and the generation of spent pickle liquor. To accom- 
plish this, a cooperative program between the Bureau of Mines and 
the American Iron and Steel Institute (AISF) was initiated. The AISI 
member companies aided research by providing materials and back- 
ground information. 

Pickling is defined here as the act of soaking in a solution for 
the purpose of cleaning or conditioning. For stainless steels, the 



' Supervisory research chemist, 
Albany, OR 



Albany Research Center, Bureau of Mines, 



surface that is eventually pickled has already been influenced by 
the annealing and scale-conditioning processes. The first process, 
annealing, determines the initial state of both the surface and of 
the chromium-depleted region. Temperature, time at temperature, 
and oxygen partial pressure during the annealing operation affects 
the thickness, structure, composition, and adhesion of the oxide 
scale to the stainless steel. All of these factors affect the ease of 
pickling and if left uncontrolled may cause a variability in the pick- 
ling process. 

The depth and composition of the chromium-depleted region 
are interrelated with the scale formation. That is, the chromium 
that's depleted from the bulk metal goes to form the oxide scale. 
Temperature, time at temperature, and oxygen partial pressure, the 
same factors that affect the annealing scale, affect the chromium- 
depleted region. This depleted region has to be removed during 
pickling to assure full corrosion resistance and it may, in fact, play 
a key role in the pickling process. 

Conditioning of the scale prior to pickling, which is not always 
done commercially, is the second process that affects the pickling 
of stainless steel. This consists usually of mechanical (shot blast- 
ing), chemical (acid), electrolytic (sulfate), or molten salt treatments 
that are aimed at facilitating the pickling process. Each of the con- 
ditioning techniques has in common the alteration of the chemical 
or physical structure of the oxide scale. While it is easy to recog- 
nize the usefulness of such conditioning techniques it may be that 
a proper understanding of the annealing and pickling processes 
would make them unnecessary. 

The actual process of pickling stainless steels is complex 
because of the environment that is usually used. The environment 



40 



consists of solutions of nitric and hydrofluoric acids at tempera- 
tures of 50° to 80° C containing large quantities of dissolved Fe, 
Cr, and Ni. To truly understand the process of pickling, and thus 
to control it, it is necessary to understand the chemistry of this com- 
plex environment. This is an environment where HF, a weak acid 
and a strong complexing agent, is combined with a strong oxidiz- 
ing acid that can be relatively easily converted into diverse other 
oxidized or reduced forms. The dissolved metals tend to tie up the 
HF, making it necessary to continually add more HF, while the 
high-temperature and chemical dissolution reaction tend to expel 
the nitric acid from solution as a nitrogen oxide. The knowledge 
that is needed consists of activity coefficients, solubilities, and sta- 
bility constants for a large number of stable and metastable com- 
pounds and complexes. 



Ingot 




Hot work 




. 


. 


„., 
































Conditioning: 

mechanical 

or 

chemical 








' 






Pickling 














Cold work 


















Anneal 














Conditioning: 
chemical 

or 
electrolytic 

or 
salt bath 














Pickling 


As necessary 
















Finished 
material 





Some smdies in most of the areas relevant to the pickling of 
stainless steels have been done. These include studies by the AISI 
members on the effects of annealing and conditioning on the pick- 
ling of stainless steels. Some basic properties of HNO3-HF solu- 
tions and methods to remove dissolved metals from these solutions 
are being studied at the Mackay School of Mines. The work to be 
reported in this paper consists of the effect of temperature, acid, 
and dissolved metal concentrations on the pickling of 304SS. 

The model of the pickling process used to design the experimen- 
tal approach consisted of an undercutting of the scale through dis- 
solution of the chromium-depleted region beneath the scale. This 
is shown schematically in figure 2. The chromium-depleted region 
on 304SS, which formed because of a short high-temperature anneal, 
has been characterized (1).'^ On the basis of this characterization, 
alloys were developed and used in the test program that is being 
reported here. Pickling was thus assumed to occur by dissolution 
of the chromium-depleted region as represented by three experimen- 
tal alloys described in the following section. Optimization of the 
pickling process would occur when the chromium depleted region 
dissolved at its fastest rate and the bulk 304SS dissolved at its slowest 
rate. Under these conditions optimization would result in the mini- 
mum amount of Fe, Cr, and Ni dissolved, the minimum amount 
of acids used, and the minimum amount of spent pickle liquor 
generated. 



2 Italic numbers in parentheses refer to items in the list of references at the end 
of this paper. 



HNO3-I-HF 




Figure 1 .—Steps used in processing stainless steel strip. 



Figure 2.— Schematic of assumed pickling process. 



41 



EXPERIMENTAL 



The commercial alloy used in all experiments was 304SS sup- 
plied by the Allegheny Ludlum Steel Corp., Leechburg, PA, and 
having the nominal composition Fe-18Cr-9Ni. The three experimen- 
tal alloys were nominally Fe-4Cr-13Ni, Fe-12Cr-12Ni, and 
Fe-12Cr-17Ni. 

All of the alloys were exposed to the HNO3-HF solutions which 
were made from reagent grade HNO3 and HF and high-purity 
(18 Mfi-cm) water. Only the 304SS was exposed to solutions con- 
taining dissolved Fe, Cr, and Ni. These solutions were made using 
reagent grade Fe(N03)3.9H20, Cr(N03)3*9H20, and 
Ni(N03)2«6H20. All solutions were deaerated using ultra-high- 
purity (UHP) nitrogen for 4 to 16 h prior to running a test. 

All weight-loss tests were conducted by exposing the sample 
for a period of time, removing from solution, rinsing with high- 



purity water, and drying with UHP nitrogen. Samples were subse- 
quently weighed to determine the weight loss, and then reexposed 
to the solution for additional periods up to a total of 90 min exposure 
time. All corrosion rates were determined by fitting the linear weight 
loss versus exposure time data to a linear equation. 

Auger electron spectroscopy (AES) measurements were taken 
using a Physical Electronics^ model 10-155 cylindrical mirror 
analyzer. The Auger electrons were induced by a 5-keV electron 
beam incident at 60° with respect to the sample normal, and a 2-V 
peak-to-peak modulation signal was applied to the analyzer. To 
obtain element depth profiles, a 2.5-keV Ar+ beam incident at 10° 
with respect to the sample normal was used to sputter away small 
increments of oxide thickness. 



RESULTS 



Data for 304SS are plotted in figures 3 through 6 to show the 
general effects of temperature, nitric, and hydrofluoric acid con- 
centrations, and dissolved iron, chromium, and nickel concentra- 
tions on the dissolution of 304SS. Buildup of dissolution products 
during the 90-min exposure had no effect on any of the dissolution 
rates reported here. This was based on the fact that the metals lost 
weight linearly with time. If dissolution products affected the dis- 
solution process then the data would tend to deviate from linearity. 
The activation energy necessary for dissolution of 304SS as calcu- 
lated from curves similar to figure 3 apj)eared generally to range 
from 5 to 10 kcal/mol in solutions with 1.3M nitric acid and from 
9 to 14 kcal/mol in solutions with lower concentrations of nitric 
acid. These categories applied also when dissolved metals were 
present. 

The data in figure 4 show that the dissolution rate of 304SS 
at constant temperature and nitric acid concentration varies approx- 
imately linearly with hydrofluoric acid concentration. This was the 
same relationship observed for the three experimental alloys. Nei- 
ther nitric acid nor dissolved iron affected this linear relationship. 
The data in figure 5 show that the dissolution rate of 304SS at con- 
stant temperature and hydrofluoric acid concentration passes though 
a maximum as a function of nitric acid concentration. This maxi- 
mum occurs at approximately 0AM to 1.5M nitric acid. 
Hydrofluoric acid concentration affected only the magnitude of the 
dissolution rate and not the shape of this curve. For the three 
experimental alloys, the shape of this curve was the same but the 
location of the maximum was shifted to higher concentrations com- 



3.5M HNO, 



3.5M HNO-,- 



2- 




-2.7M HF 



0.0030 
RECIPROCAL TEMPERATURE, K 



pared to that for 304SS and the overall height of the curve increased 
with decreasing chromium concentration. 

A typical pickling solution usually has large quantities of dis- 
solved iron, chromium, and nickel, but because of the composi- 
tion of the stainless steels the dissolved iron is always more 
concentrated than either dissolved chromium or nickel. A compar- 
ison of the effect of dissolved iron, chromium, and nickel on the 
dissolution rate of 304SS in HNO3-HF solutions is shown in figure 
6. Both iron and chromium cause a decrease in dissolution rate. 
Iron was the most effective in reducing the dissolution rate of the 
304SS, dissolved chromium was less effective than iron, while dis- 
solved nickel had no apparent effect. Iron was also the only dis- 
solved metal species to shift the intercept of the curve. 

Films formed as a result of exposure to HNO3-HF were ana- 
lyzed using AES. Profiles were obtained by alternately doing AES 
analyses and argon ion sputtering at a rate of 1 .5 A min (calibrated 
using Ta205). Profiles of 304SS samples exposed to different 
HNO3-HF solutions show an enrichment in Cr in the film com- 
pared to the metal substrate, and a small quantity of fluorine in the 
region of the film-environment interface. The fluorine level was 



' Reference to specific products does not imply endorsement by the Bureau of Mines. 



E -^ 



< 

a: A 



- 




/^8M HNO3 


- 


♦/ 


^^^ 3.5M HNO3 


■^ 


1 


O.OM HNO3 
1 



Figure 3.— Arrhenius plots for the dissolution of 304SS in 
HNO3-HF solutions. 



12 3 

HF CONCENTRATION, mol/L 

Figure 4.— Effect of HF concentration on dissolution rate of 
304SS in HNOa-HF solutions at 50« C. 



42 



^ 1.00 
u 

c 

~^ .75 
E 



.50 



- o 


^ 












- / 


\ 


\ ^ 










- 1 








o 






" 1 




D 










- 1 y^ 


' □ 








^^ c 




- /S 






^~~~~~ 


---~-__7o; c 




o 


1/ 








n 

sn" r 


— — 


— — ~s_ 


(Lh — PT 


A 




30- 


c 


s 


^ — - 


P, , 


1 1 


, 1^, ,M 


1 1 1 1 


1 1 1 1 



12 3 4 5 

HNO3 CONCENTRATION, mol/L 

Figure 5.— Effect of temperature and HNOa concentration on 
dissolution rate of 304SS in HNO3-HF solutions with approxi- 
mately 0.5M HP concentration. 



3.5M HNOj, 70'C 

No metal added 




.2M Fe - 



12 3 4 

HF CONCENTRATION, mol/L 

Figure 6.— Effect of equimolar concentrations of dissolved Fe, 
Cr, and Ni on dissolution rate of 304SS in HNO3-HF solutions at 
TO' C containing approximately 3.5Af HNOa. 



independent of HF concentrations between 0.5M and 1.5M HF. 
There was some evidence that the level of fluorine was dependent 
on the solution at concentrations below 0.5M HF. The main differ- 
ence was a variation in film thickness by about a factor of 2 between 
the thickest and thinnest films. This variation in thickness was not 
correlated with any of the test parameters such as temperature or 
concentration. 

The fluorine profile in figure 7 shows that a fraction of a 
monolayer of fluorine exists on the oxide surface below the carbon 
surface contamination. The surface carbon consists primarily of 
hydrocarbons picked up from exposure to the atmosphere follow- 
ing the HNO3-HF treatment. The conclusion that the fluorine exists 
on the oxide surface is based on many profiles identical to that shown 
in figure 7. There was a general correlation between the time 
required to sputter away the carbon surface contamination and the 
fluorine. It appeared that the fluorine was rapidly removed only 
after most of the surface carbon was removed. The relatively flat 
portions of the fluorine profiles represent regions where only sur- 
face contamination was being removed. If all of the fluorine was 
concentrated at the film surface (i.e., at the film-carbon contami- 
nation interface), then the fluorine could represent as much as one- 
quarter to one-half a monolayer of fluorine at that surface. 




SPUTTERING TIME 



Figure 7.— AES sputter depth profile of film formed on 304SS 
exposed to 3.5Af HNOs-O.SAf HF-0.2Af Fe solution for 90 min at 
70" C. The sputter rate was approximately 1 .5 A min. 



DISCUSSION 



The first part of this discussion will be applied to understand- 
ing the mechanism of dissolution of metals in HNO3-HF solutions. 
The second part will concentrate on applying the data to optimiz- 
ing the pickling process. 

The range of activation energies (5-14 kcal/mol) measured here 
are similar to those reported (2-8) for the dissolution of rapidly 
corroding metals in solutions other than HNO3-HF. This is in con- 
trast to those activation energies (15-22 kcal/mol) reported (7-8) 
for the dissolution of metals that passivate. It therefore appears that 
304SS in the HNO3-HF solution is not inhibited from dissolving 
by a passive film. 

The evidence obtained from the AES measurements suggests 
that there is a film on the surface of the metal in solution and that 
this film is not a passive or protective type of film. Research (9) 



done elsewhere on 304SS in similar solutions supports the conclu- 
sion that this is not a passive film. A truly passive film in HNO3 
solutions would exhibit a more intense chromium peak than that 
shown in figure 7. The presence of the fluoride on the outer sur- 
face of this film suggests that fluoride is intimately involved in the 
dissolution mechanism. This is supported also by the observed lin- 
ear increase in dissolution rate with increasing HF concentration. 
The dissolution rate vs HF curves for HNO3-HF solutions were 
extrapolated to zero dissolution rate at an HF concentration of zero. 
This is in agreement with the reported (10) zero dissolution rate 
of 304SS in low-temperature HNO3. The HN03-HF-Fe solutions 
do not, however, pass through zero dissolution rate at a zero con- 
centration of HF. This is probably related to the formation of iron- 
fluoride complexes. These complexes are reported (11) to be very 



43 



stable and would be able to tie up the available fluoride. The inter- 
cept of the curves in figure 6 corresponds to approximately a 3:1 
molar ratio of F to Fe, which suggests the formation of FeFj 

The iron, chromium, and nickel additions to the HNO3-HF solu- 
tions cause various degrees of reduction of the dissolution rate of 
the 304SS. Dissolved iron has the greatest inhibiting effect, fol- 
lowed by dissolved chromium, which is less effective, and dissolved 
nickel, which has little effect on the dissolution rate. The effec- 
tiveness of these metal ions in inhibiting the dissolution of 304SS 
appears to be due to the specific cation's ability to tie up the avail- 
able fluoride anions. This can be seen by considering the stability 
constants for the metal-fluoride ligand involving one metal and one 
fluoride ion. The stability constants (12) are 1.5x10^ for FeF^*, 
2.3 X 10" for CrF2^ and 6.3 for NiF^ Thus, the order of effective- 
ness of the metal ions is the same as the order of the stability cons- 
tants. The negligible effect of the dissolved nickel is reasonable 
because the stability constant (II) of 6.3 for NiF^ is very similar 
to the value of 3.9 for the most prevalent (12) ionized fluorine 
species, HF2. 

The role played by dissolved metals in reducing the dissolu- 
tion rate of 304SS appears to be one of reducing the amount of fluo- 
ride available for participation in the dissolution reaction. These 
dissolved metals do not participate directly in the dissolution reac- 
tion, but rather they alter the solution chemistry. This emphasizes 
the importance of the fluoride component of the pickling bath. The 
finding, by AES measurements, of the fluoride on the surface of 
the metal's film becomes a key to the mechanism of dissolution 
of the metal. Others (13-14) have suggested that HF has a cata- 
lytic effect in the formation of metal-fluoride complexes at the film- 
solution interface. It is postulated that HF allows a more rapid trans- 
fer of these complexes into solution as compared to species formed 
in non-fluoride-containing solutions. 

The data in figure 5 provide evidence of the catalytic nature 
of the dissolution reaction. The shape of the dissolution rate versus 
nitric acid concentration curve is characteristic of a reaction that 
proceeds through heterogeneous catalysis. In such a case the cata- 
lyst typically adsorbs to the surface where the reaction occurs and 
the catalyst remains unchanged. For the dissolution of 304SS in 
HNO3-HF solutions it appears that the fluoride ion adsorbs to the 
nearly passive film formed in solution, enhances the dissolution of 
this film, and then complexes with the dissolved metal species. For 
this mechanism to continue it is necessary for the film to continu- 
ally reform and this can more than adequately be done by the reac- 
tion of the nitric acid with the metal. 

The data shown in figure 5 make it possible to consider optimiz- 
ing the pickling process. The maximum and minimum represent 
two areas where dissolution can be very fast or very slow. It was 
already noted that a similar type of behavior was observed for the 
three alloys representative of the chromium-depleted region. Thus 
to optimize the pickling process in terms of HNO3 concentration 
it is only necessary to superimpose the graphs for 304SS and one 
or more of the experimental alloys. This has been done schemati- 
cally in figure 8 for 304SS and the Fe-12Cr-17Ni alloy. The other 
alloys had basically the same shape but a greater height. The region 
on figure 8 labeled as optimum represents the range of nitric acid 
concentrations over which the pickling of 304SS should be 
opfimized. At 50° C that range is 0.8M to 1.3M HNO3 in solu- 
tions with 0.5Af HF. It is in this range that the dissolution of the 
chromium-depleted region will be maximized and the dissolution 
of bulk 304SS will be minimized. It can be seen that being to the 
left of this region would sacrifice base metal for dissolution (pick- 
ling) speed while being to the right will result in a very slow pickling. 

The other factors to consider are temperature and HF concen- 
tration. The HF concentration should linearly increase pickling rate 



and loss of base metal so that it is best to use as low a concentra- 
tion as practical. Choice of temperature is another variable that is 
determined somewhat by practicality. For the activation energies 
reported here, there is an order of magnitude increase in dissolu- 
tion (pickling) rate of 304SS for approximately each 20 ° to 40 ° 
C rise in temperature. 

The dissolution rate data can also be used to develop an equa- 
tion describing the overall response of dissolution rate to time, tem- 
perature, and acid concentrations. The equation would take the form 
of a heterogeneous catalysis rate equation and be similar to the fol- 
lowing equation: 



Dissolution rate = 



k [HFp [HN03]« 
1 -I- [HF]" [HN03]= 



exp (-AH/RT), 



where k, a, B, and R are constants, AH is the activation energy, 
T is the absolute temperature, and [HF] and [HNO3] are acid molar- 
ities. Such as equation will help to give a better insight into the 
dissolution mechanism (depending on the fitted values of a, B, and 
AH) and can also be used to predict dissolution (pickling) rates 
for acid concentrations and temperatures between those actually 
measured. 

Equations such as that described for 304SS and for the three 
experimental alloys can be used further to model the amount of metal 
lost during the pickling reaction. This would be accomplished in 
terms of the overall basic assumption made in this paper. That is, 
that pickling occurs by a dissolution of the chromium-depleted 
region, undercutting the scale and thus removing it. The modeling 
would proceed by assuming a thickness of chromium depletion, an 
amount of scale formed during annealing, a preset pickling time, 
and the number of pickling-working-annealing steps in the life of 
a coil of stainless steel. The result would be the minimum amount 
of time needed to remove the chromium-depleted region (and thus 
the scale) and the minimum amount of metal lost (and metal lost 
unnecessarily). 

There are certain quantities that would make this predictive 
model more accurate. The quantities that are relatively unknown 
at present are represented by the chromium-depleted region and 
the annealing scale. The needed information would be in the form 
of equations expressing the effect of annealing time, temperature, 
and oxygen content of the environment on both the chromium- 
depleted region and the oxide scale composition and thickness. These 
would be rewarding areas in which to conduct research. 




HNO3 CONCENTRATION ► 

Figure 8.— Schematic of optimum range of HNO3 concentra- 
tions for piclding 304SS in HNO3-HF solutions. 



44 



FUTURE WORK 



Additional work in this area at the Bureau's Albany (OR) 
Research Center will consist solely of data analysis. Data analysis 
will be completed and equations developed to model the dissolu- 
tion rate as a function of acid concentrations and temperature. These 
will then be used to model the loss of metals during pickling with 



the hope of being able to best predict optimum conditions for pick- 
ling stainless steels. This will also include completing the analysis 
of work on 430SS and experimental alloys used to represent its chro- 
mium depleted region. 



CONCLUSIONS 



1. Increasing hydrofluoric acid concentration causes a linear 
increase in the dissolution rate of 304SS. 

2. Increasing nitric acid concentration up to 0AM to 1.5M 
HNO3 causes an increase in the dissolution rate of 304SS, and a 
decrease in the dissolution rate for higher HNO3 concentrations. 

3. Dissolved iron and chromium reduce the dissolution rate of 
304SS in HNO3-HF solutions. Dissolved nickel has little or no 
effect. 



4. The surface films present on 304SS in HNO3-HF solutions 
did not change as a function of dissolution rate and contained a frac- 
tion of a monolayer of fluoride on the outer surfaces of the films. 

5. Optimum pickling conditions exist at nitric acid concentra- 
tions of O.SMto I.3MHNO3 at 50° C in solutions with 0.5MHF. 



REFERENCES 



1. Fabis, P. M., and B. S. Covino, Jr. Near Surface Elemental Con- 
tration Gradients in Annealed 304 Stainless Steel as Determined by Ana- 
lytical Electron Microscopy. Oxid. Met., v. 25, Nos. 5/6, 1986, pp. 
397-407. 

2. Muralidharan. V. S.. and K. S. Rajagopalan. Kinetics and Mecha- 
nism of Corrosion of Iron in Phosphoric Acid. Corros. Sci., v. 19, 1979, 
pp. 199-207. 

3. Alexander, B. J., and R. T. Foley. Anion Dependence of the Acti- 
vation Energy for Iron Corrosion. Corrosion, v. 31, No. 4, 1975, pp. 
148-149. 

4. Altura, D., and K. Nobe. Activation Energy for the Corrosion of 
Iron in Sulfuric Acid. Corrosion, v. 20, No. 11, 1973, pp. 433-434. 

5. Makrides, A. C, and N. Hackerman. Solution of Metals in Aque- 
ous Acid Solutions. II Depolarized Solution of Mild Steel. J. Electrochem. 
Soc, V. 105, 1958, pp. 156-162. 

6. Riggs, O. L. Activation Energy from Carbon Steel Corrosion in Phos- 
phoric Acid. Corrosion, v. 24, No. 5, 1968, pp. 125-126. 

7. Covino, B. S., Jr., J. P. Carter, and S. D. Cramer. The Corrosion 
Behavior of Niobium in Hydrochloric Acid Solutions. Corrosion, v. 36, 
No. 10, 1980, pp. 554-558. 



8. Ishikawa, T., and G. Okamoto. Potentiostatic Response of Passive 
Metals to the Rate of Temperature Change. Electrochimica Acta, v. 9, 1964, 
pp. 1259-1268. 

9. Asami, K., and K. Hashimoto. An X-ray Photoelectron Spectroscopic 
Study of Surface Treatments of Stainless Steels. Corros. Sci., v. 19, 1979, 
p. 1007. 

10. Wilding, M. W., and B. E. Paige. Survey on Corrosion of Metals 
and Alloys in Solutions Containing Nitric Acid. Allied Chemical Corp., 
ICP-1107, 1976, 56 pp. 

11. Smith, R. H., and A. E. Martell. Critical Stability Constants, Vol- 
ume 4— Inorganic Complexes. Plenum (New York), 1976, pp. 96-103. 

12. Pourbaix, M. Atlas of Electrochemical Equilibria in Aqueous Solu- 
tions. Pergamon (New York), 1966, p. 587. 

13. Lochel, B., and H. H. Strehblow. Breakdown of Passivity of Iron 
by Fluoride. Electrochim. Acta, v. 28, No. 4, 1983, pp. 565-571. 

14. Lochel, B. P., and H. H. Strehblow. Breakdown of Passivity of Nickel 
by Fluoride. II Surface Analytical Studies. J. Electrochem. Soc, v. 131, 
1984, p. 713. 



45 



DECREASED ACID CONSUMPTION IN STAINLESS STEEL PICKLING 

THROUGH ACID RECOVERY 



By G. L. HorteM and J. B. Stephenson^ 



ABSTRACT 



Acid pickling of stainless steel annually generates approximately 30 million gal of spent nitric 
acid-hydrofluoric acid solutions as a byproduct. Disposal of these acid solutions significantly increases 
the cost of manufacturing stainless steel and results in the loss of the acid, Cr, and Ni values. Recent 
advances in ion-selective membrane technology have opened new avenues to regenerate these acid 
solutions as an alternative to disposal. Experimental work by the Bureau of Mines has indicated 
that an electrodialysis cell utilizing ion-selective membranes has potential for separating dissolved 
metals from spent pickling acid solutions, while regenerating the acids for return to the pickling 
process. 



INTRODUCTION 



The Bureau of Mines is conducting research to conserve mineral 
values and reduce waste generation from spent nitric acid- 
hydrofluoric acid pickling solutions. The stainless steel industry 
annually generates approximately 30 million gal of these solutions, 
which currentiy have to be neutralized and sent to waste disposal 
(4).' 

Hot working and annealing of stainless steels cause a strati- 
fied scale to form on the surface of the steel . The scale consists 
of metal oxides in the top layer, an intermediate layer of spinel oxide 
[FeO(Fe(2-x)CrJ03, where 0<x<2)], and a layer of steel, 
which has been depleted of Cr and Ni. The concentration of Cr 
and Ni in the oxide scales may be four to six times their respective 
concentrations in the bulk alloy (2-3, 16). 

Scale removal relies on a variety of conditioning and pickling 
steps (1). The commonly used acid pickling solutions contain 5 to 
25 vol pet nitric acid and 0.5 to 8 vol pet hydrofluoric acid (1). 



' Chemical engineer. 
^ Research chemist. 

Rolla Research Center, Bureau of Mines. Rolla, MO. 
' Itaiic numbers in parentheses refer to items in the list of references at the end 
of this paper. 



Evidence suggests that the chemical pickling solution does not dis- 
solve the oxide scale but dissolves the underlying metal, allowing 
the scale to fall away (15). 

The pickling acids become ineffective as the concentration of 
dissolved Fe increases, after which the solutions are discarded. This 
necessitates the disposal of the pickling solutions. Disposal according 
to the required environmental standards has been estimated to 
increase the total operating cost of the steel industry by 8 to 10 pet. 

Several processes have been developed for reclaiming HNO3 
and HF from pickling soludons but the complexity of these methods 
has limited their acceptance by domestic steel producers. Most nota- 
ble of the available reclamation processes rely on distillation- 
crystallization or solvent extraction technology (14). 

Preliminary Bureau research focused on adapting a Bureau- 
developed (6, 11-12) electrolytic process for regenerating chro- 
mic acid solutions. Electrolytic processing of stainless steel pick- 
ling solutions would be attractive because initial capital costs are 
generally lower than for solvent extraction or chemical process- 
ing. In addition, energy requirements are usually less for electro- 
lytic processes than for distillation or evaporation processes. This 
paper reports on research in progress on the electroregeneration 
of spent stainless steel pickling solutions in a Bureau-developed 
three-compartment membrane cell. 



46 



BACKGROUND 



PICKLING CHEMISTRY 



The oxide scale fonned during hot working and annealing of 
stainless steel is enriched in Cr and Ni. Chromium and nickel dif- 
fuse from the bulk steel to the oxides, changing the alloy composi- 
tion at the steel surface to a high Fe, low Cr-Ni alloy. The Cr content 
in this surface alloy is insufficient to provide a protective, passive 
oxide film for adequate corrosion resistance. To regain suitable cor- 
rosion resistance, the oxides and Cr-deficient surface alloy must 
be removed from the steel. 

Heavy oxide scale produced during hot rolling typically is 
treated and removed prior to acid pickling by shot blasting. Cold- 
rolled and annealed material have thinner oxide layers. However, 
in both situations, the Fe-enriched surface alloy is the principal mate- 
rial dissolved by acid pickling. Iron is the predominant metal in 
solution because of this enrichment, as shown in table 1 (1). 

Table 1. — Typical concentrations In spent pickling solutions 



Component 


Concentration 




g/L M 


Nitrate 

Fluoride 

Iron 

Chromium 


144 2.32 

46 2.42 

34 .61 

6 9.34 


Nickel 


6 .10 



Literature review indicates that mechanisms and kinetics for 
metal dissolution in this acid system are complex and not fully under- 
stood. Research has shown that both HNO3 and HF play a role in 
dissolving metal (2). It is believed that nitric acid reacts with the 
surface to convert metals to oxides. This is consistent with the 
oxidizing behavior of nitric acid. Hydrofluoric acid reacts with the 
oxides to form soluble fluoride salts. Fluoride is a strong ligand, 
which readily complexes Fe in solution. The equilibria relation- 
ships for iron fluoride salts, given in table 2, indicate that an iron 
fluoride salt should be the principal metal species in solution. The 
relative abundance of fluoride in relation to Fe in typical spent pick- 
ling solutions indicates that the predominant specie would be the 
cation FeF2+ (7, 9, 13). 



considered. An applied electromotive force provides a strong sepa- 
ration driving force by attracting cations to the cathode. Metal 
removal by electrolytic deposition from the pickling solutions would 
not be expected because of the presence of HNO3; therefore, a 
receiving solution should be provided as a sink for the dissolved 
metals. HNO3 reacts with the metallic anode to liberate NO^; 
therefore, it would be desirable to avoid contacting the pickling 
solutions with the anode. Hydrogen ions generated as an electroly- 
sis product at the anode must be allowed to migrate into the pick- 
ling solution in order to regenerate nitric and hydrofluoric acids. 
Physically separating the solutions in an electrolysis cell, therefore, 
requires the use of permeable barriers that allow transport of the 
dissolved metals and hydrogen ions, while retaining the acids in 
the pickling solution. 

Perfluorinated sulfonic acid (PFSA) cation exchange membrane 
was the most suitable permeable barrier commercially available 
when the research was initiated. The PFSA membranes are pre- 
pared from long-chain fluorinated polymers with material charac- 
teristics similar to Teflon 'fluorocarbon polymer. Sulfonic acid 
groups incorporated in the polymer matrix act as exchange sites, 
allowing migration of cations across the membrane while rejecting 
anions (5). 

The type of electrolysis cell incorporating all the above require- 
ments was an electrodialysis cell (fig. 1). The PFSA membranes 
separate the pickling solution from the electrodes, the anolyte pro- 
vides a source of hydrogen ions, and the receiving solution (catho- 
lyte) provides a sink for the dissolved metals. 

The movement of ions in the electrodialysis cell is regulated 
by the PFSA membranes and the current flux in the cell. The 
charged ions attempt to move to the oppositely charged electrode. 
The anions (A~) are attracted to the positively charged anode; 
however, movement is blocked by the PFSA membrane. There- 
fore, the majority of anions should be retained in the pickling solu- 
tion. The cations (C+) are attracted to the negatively charged 
cathode and are accepted for transport by the PFSA membrane. 
Hydrogen ions are produced at the anode and migrate to the cath- 
ode. As the hydrogen ions pass through the central compartment, 
they will replace the metal cations that have been removed, gener- 
ating HNO3 and HF. However, part of the hydrogen ions gener- 
ated at the anode will pass through the pickling solution to the 
cathode where they are reduced to hydrogen gas. 



Reference to specific products does not imply endorsement by the Bureau of Mines. 



Table 2.— Formation constants for iron fluoride salts (9) 



Fe3+ 


+ F- 


- FeF2 


Fe3-^ 


+ 2F- 


^ FeFz 


Fe3+ 


+ 3F- 


- FeF3 


Fe3 + 


+ 4F- 


^ FeFj 


Fe3+ 


+ 5F- 


-* FeFs 



1.58.105 
1.26. 109 
1.00. 10'2 
1.00.101" 
3.16.101" 



Increasing the hydrofluoric acid content of the bath results in 
speciation changes toward higher fluoride salts. A solubility limit 
for FeFj and CrFj salts occurs at a total metal content of approxi- 
mately 50 g/L. Adding HF to a solution at this metals content dra- 
matically increases the danger of salt crystallization. Therefore, the 
pickling solution is usually discarded before the 50-g/L metal con- 
tent is reached. 



THE ELECTRODIALYSIS PROCESS 

Before applying an electrolysis process to HNO3-HF pickling 
solutions, the possible electrode reactions that can occur must be 



(-) 



Catholyte Pickle 

solution 



Cations 



6 Anions 



Anolyte 



(+) 



Excess H 




— H 
Membrane 



Cathode Membrane 

Figure 1 .—Electrodialysis cell 



Anode 



47 



EXPERIMENTAL METHOD AND RESULTS 



LABORATORY CELL DESIGN 

The laboratory cell design utilized a plate and frame construc- 
tion to allow flexibility in size and configuration of the cells (8). 
The design was based on use of U-shaped frames to form the liquid- 
holding compartments and plates on the outer ends to seal the com- 
partments. Capacity of the cells ranged from 300 mL to 1 L per 
compartment. The working dimensions of each compartment were 
10.2 cm wide by 27.9 cm high by 1 .3 cm deep for the small frames 
and 5.0 cm deep for the large frames. A seal between the frames 
was obtained with neoprene gaskets. The frames were held together 
by stainless steel bolts placed on 2.5-cm centers around the perim- 
eter of the cell. A 316 stainless steel cathode and a Pb-Sb anode 
were used; each with working dimensions of 9.21 cm by 25.4 cm 
for a working area of 2.34 dm^. The membrane window was 10.2 
cm by 25.4 cm for an effective area of 2.59 dm^. Circulation was 
provided by means of an input pump and a simple overflow incor- 
porated into the compartment outer wall. 



CELL PERFORMANCE 

The effectiveness of the electrodialysis cell was initially deter- 
mined on two synthetic pickling solutions. A statistically designed 
experiment utilizing Plackett-Burman screening methodology was 
performed as an initial investigation of the electrodialysis regener- 
ation process (10). This method involves selecting maximum and 
minimum values for the system parameters of interest and perform- 
ing a series of trials using combinations of the parameter values. 
Statistical analysis of the system responses indicates which of the 
parameters had an effect on the system performance. 

The system parameters selected for study were the form of the 
dissolved metal complex in either a high HNO3 or a high HF solu- 
tion, H2SO4 concentration in the anolyte and catholyte, and the 
applied current density. The maximum and minimum values for 
these parameters were selected according to available industrial data 
and the physical constraints of the system (table 3) (8). 



The fluoride and nitrate levels in the pickling solution were 
varied to control the complex species present and determine its effect 
on Fe transport across the membrane. Iron complexes with fluo- 
ride in preference to nitrate, therefore, adjusting the abundance of 
fluoride in the system ensured the predominance of either monova- 
lent or divalent iron fluoride salts (table 4). The HF-based stock 
should contain predominantly FeF^+ in equilibrium with the nitrate 
in solution. The HNOs-based stock should contain predominantly 
FeF2+. 

The selected system responses were the quantity of Fe, F, and 
nitrate extracted from the pickling solution. To regenerate the pick- 
ling solution, the Fe must be extracted in greater quantity than the 
nitrate and fluoride. Failure to achieve this would indicate that the 
electrodialysis would have little potential for regenerating the pick- 
ling solution. 



Table 3. — Parameters used in Plackett-Burman designed 
experiment 



Parameter 



Minimum 



Maximum 



Pickle type, 1/W 

Concentration, vol pet H2SO4: 

Catholyte 

Anolyte 

Current density. . . .A/dm^ . . 



FeFa, HNO3 


Fe(N03)3, HF 


1 


10 


1 


10 


1.1 


5.4 



Table 4.— Pickling solution analyses 



Pickle 


Molar concentrations 


Molar ratios 


type 


Fe 


F 


NO3 


F-Fe 


N03-Fe 


HF-based 

HNOj-based . . . 


0.82 
.92 


1.04 
2.90 


2.45 
1.10 


1.3:1 
3.2:1 


3.0:1 
1.2:1 



DISCUSSION 



Statistical analysis of the system responses showed a response 
dependency on two of the four varied parameters (table 5) . Note 
that the type of pickling solution had no effect on Fe extractions 
but did affect extractions of fluoride and nitrate. Current density 
had a significant effect on iron extraction since it was related to 
the driving force for ion transport in the cell. The PFSA membrane 
rejects anions: therefore, the dependency of fluoride extraction on 
current density indicates transport of a cationic iron-fluoride com- 
plex. It is significant that anolyte and catholyte concentrations had 





Table 5. 


—Parameter effects 




Ion 


Pickle 
type 


Concentration 


Current 


extracted 


Catholyte 


Anolyte 


density 


Iron 

Fluoride . . . 
Nitrate .... 


No... 
Yes... 
Yes . . . 


No 

No 

No 


No . .. 
No . . . 
No ... 


Yes. 
Yes. 
No. 



no effect on Fe extraction. Iron transport independent of the sul- 
furic acid concentration in the anolyte and catholyte allows greater 
flexibility for optimizing the efficiency of the cell. 

The extractions achieved during the experiments indicate that 
current density had the most notable effect on Fe extraction (table 
6). Note that the fluoride extraction was also greatest for the experi- 
ments performed at high current density. The difference in nitrate 





Table 6.— Extractions 




Parameter 


Runs Extraction, pet 


Molar ratio 




Fe F NO3 


F-Fe N03-Fe 



Average 


12 


35.09 


26.46 


11.35 


HF-based .... 


6 


34.09 


28.38 


15.43 


HN03-based . . 


6 


36.01 


24.54 


7.26 


5.4 A/dm2 


6 


52.99 


37.37 


12.76 


1.1 A/dm2 


6 


17.68 


11.73 


9.94 



1.7:1 


0.7 


1.1:1 


1.4 


2.2:1 


.2 


1.6:1 


.5 


1.5:1 


1.2 



48 



extractions for the 5.4-A/dm^ runs and the l.l-A/dm^ runs may 
be significant, since the higher current density represents a stronger 
driving force; however, the statistical analysis of the responses (table 
5) indicated no relationship between current density and nitrate 
extraction. 

Comparing the molar ratios of the anions to Fe in the extracted 
fractions provides a clue as to the mechanism of anion extraction. 
The molar ratios for the HNOa-based pickling solution indicated 
that FeF2+ was the predominant specie transported across the 
membrane (fig. 2). Transport of nitrate from the HNOs-based pick- 
ling solution was not indicated. The experimental runs with the HF- 
based pickling solution showed that fluoride was extracted unimo- 
lar with Fe, indicating that FeF^+was the predominant specie trans- 
ported. The fluoride extractions from both pickling solution types 
were consistent with the equilibria constants for iron-fluoride com- 
plexes. 

The current density had a dramatic effect on the extraction of 
Fe. This was due to the electromotive driving force in the cell, the 
higher current density representing the higher flux and hence the 
strongest driving force for transport of Fe. The range of Fe trans- 
port rates during the test series was 0.0093 to 0.0372 g-mol/dm^h, 
where the area term relates to membrane area (7). 

The variation in current density had little effect on transport 
of fluoride relative to Fe, supporting the theory that the fluoride 
was transported with the Fe as a complex species. The extraction 
of nitrate was almost equivalent for both current density levels 
indicating that loss of nitrate was only slightly affected by current 
density, which is consistent with table 5. This indicates that the 
cationic membrane is effective but does not totally reject transport 
of nitrate ions. 

The desired separation should remove Fe while retaining nitrate 
and fluoride in the regenerated solution. Comparison of the anion 
to Fe ratios for the unprocessed pickling solutions (see table 4) and 
the retained species in the processed pickling solutions (table 7) 
indicates that the separation is occurring. An increase in the molar 
ratios of anions to Fe after processing would indicate that Fe was 
being separated from the nitrate and fluoride (fig. 3). 

Except for the experimental runs at 1 . 1 A/dm^, the molar ratios 
indicated that nitrate and fluoride were being concentrated with 
respect to Fe; that is, more fluoride and nitrate were being retained 

Table 7.— Retention in pickling solution 



Parameter 


Runs 


Retention, 


pet 


Molar ratio 




Fe 


F 


NO3 


F-Fe 


NOs-Fe 


Average 


12 


64,91 


73.94 


88.65 


2.6:1 


2.8 




HF-based .... 


6 


65.91 


71.62 


84.57 


1 .4:1 


3.8 




HNOa-based. . 


6 


63.91 


75.46 


92.74 


3.8:1 


1.7 




5.4 Aydm2 


6 


47.02 


62.64 


87.25 


3.0:1 


3.8 




1.1 A/dm2. . . . 


6 


82.32 


88.26 


90.07 


2.4:1 


2.2 





in the pickling solutions than Fe. The results for the 1.1-A/dm^ 
runs showed no significant change from the initial solution, indicat- 
ing that the chemical driving force of the system was stronger than 
the electromotive driving force in this case. 



u- 2 



o 



ai 
cr 



en 




Average 



J3)3 FeF3 5.4 n/dm2 1.1 fl/dm^ 



Pickle type Current density 
Figure 2. — Fluoride-to-iron ratios in extracted fractions. 



c 
o 



o 



CE 

cr 



cr 
cc 




Average Average 5.4 fl/dm^ 1.1 fl/dm^ 
Stock I Processed 1 

Figure 3.— Anion-to-iron ratios in stock and processed solutions. 



49 



CONCLUSIONS 



Results of the series of experiments indicate that Fe can be sepa- 
rated from nitrates and fluorides in spent stainless steel pickling 
solutions using an electrodialysis cell. Extraction of fluorides from 
the pickling solutions show that a cationic iron fluoride complex 
was the most likely specie being transported across the membrane. 
FeF2+ and FeF2+ were the inferred species. Current density had 
a significant effect on Fe extraction, which indicates that current 
density is a controlling parameter for the process. Iron transfer rates 
ranged from 0.0093 to 0.0372 g-mol/dm^-h. PFSA membranes 
withstood exposure to the highly corrosive pickling solutions without 
a significant loss of selectivity. 



Nitrate losses from the pickling solutions indicate that the PFSA 
membrane does not completely reject nitrate transport. However, 
the nitrate extractions experienced during these tests were not 
believed to be detrimental. The overall results indicate the elec- 
trodialysis cell can transfer Fe from the pickling solutions and 
regenerate HNO3 and HF in the pickling solutions. Development 
of a technique to control the fluoride transfer may result in a via- 
ble regeneration process. 



REFERENCES 



1. American Society for Testing and Materials. Cleaning and Descal- 
ing Stainless Steel Parts, Equipment, and Systems. A380-78 in 1982 Annual 
Book of ASTM Standards; Part 3, Steel-Plates, Sheet, Strip, Wire; Metal- 
lic Coated Products; Fences. Philadelphia, PA, 1982, pp. 274-276. 

2. Bombara, G. Potentiostatic Anodic Pickling of Stainless Steel. J. Elec- 
trochem. See, v. 118, No. 4, Apr. 1971, pp. 676-681. 

3. Clark, F. H. Metals at High Temperatures. Reinhold Publ. Corp., 
New York, 1950, 372 pp. 

4. Desy, D. H. Iron and Steel. Ch. in Mineral Facts and Problems, 
1980 Edition. BuMines B 671, 1981, pp. 455-480. 

5. Eisenberg, A., and H. L. Yeager. Transport Properties of Per- 
fluorosulfonate Polymer Membranes. Ch. in Perfluorinated lonomer Mem- 
branes. ACS, 1982, pp. 41-63. 

6. Horter, G. L., and L. C. George. Demonstration of Technology To 
Recycle Chromic Acid Etchants at Gould, Inc. Paper in Proceedings, 4th 
Recycling World Congress and Exposition, New Orleans, LA, Apr. 5-7, 
1982, pp. M/3/5/1-M/3/5/13. 

7. Horter, G. L., and J. B. Stephenson. Recycling Stainless Steel Pickle 
Liquors by Metathesis With an Electromembrane Cell. Pres. at 4Ist Ann. 
Purdue Ind. Waste Conf., West Lafayette, IN, May 13-15, 1986, 16 pp; 
available from G. L. Horter, BuMines, Rolla, MO. 

8. Horter, G. L., J. B. Stephenson, and W. M. Dressel. Permselective 
Membrane Research for Stainless Steel Pickle Liquors. Paper in Proceed- 
ings of the International Symposium on Recycle and Secondary Recovery 



of Metals (Ft. Lauderdale, FL, Dec. 1^, 1985). Metall. Soc. AIME, 1985, 
pp. 467-475. 

9. Kragten, J. Atlas of Metal-Ligand Equilibria in Aqueous Solution. 
Halsted Press, 1978, 779 pp. 

10. Plackett, R. L., and J. P. Burman. The Design of Optimum Mul- 
tifactorial Experiments. Biometrika, v. 33, 1946, pp. 305-325. 

11. Soboroff, D. M., J. D. Troyer, and A. A. Cochran. Regeneration 
of Waste Metallurgical Process Liquor. U.S. Pat. 4,337,129, June 29, 1982. 

12. Spotts, D. A. Economic Evaluation of a Process To Regenerate Waste 
Chromic Acid-Sulfuric Acid Etchants. BuMines IC 8931, 1983, 7 pp. 

13. Stephenson, J. B., G. L. Horter, and H. H. Dewing. Iron Removal 
and the Complexity of Stainless Steel Pickling Liquors. Ch. in Iron Con- 
trol in Hydrometallurgy, ed. by J. E. Dutrizac and A. J. Monhemius. Hal- 
stead Press, Div. Wiley (New York), 1986, pp. 571-581. 

14. Stephenson, J. B., R. S. Kaplan, and J. C. Hogan. Recycling and 
Metal Recovery Technology for Stainless Steel Pickle Liquors. Environ. 
Prog., v. 3, No. 1, Feb. 1984, pp. 50-53. 

15. Vicentini, B., and G. Bombara. On the Mechanism of Scale Removal 
in the Acid Pickling of Austenitic Stainless Steels. Electrochim. Metall., 
v. 3, 1968, p. 313. 

16. Whittle, D. P., and G. C. Wood. Complex Scale Formation on an 
Iron— 18% Chromium Alloy. J. Electrochem. Soc, v. 114, No. 10, Oct. 
1967, pp. 986-993. 



50 



RECYCLING OF STAINLESS STEELMAKING DUSTS AND OTHER WASTES 



By L. A. NeumeieM and M. J. Adam^ 



ABSTRACT 

There are significant amounts of Cr, Ni, and other metals contained in furnace dusts, mill 
scale, and swarfs produced as wastes annually by the domestic specialty steelmaking industry. Bureau 
of Mines research has led to development of a process for the in-plant recovery of about 90 pet 
of the Cr and Mo and well over 90 pet of the Ni and Fe from stainless steelmaking electric furnace 
dusts, argon-oxygen decarburization (AOD) vessel dusts, mill scale, and oily swarf. In the proc- 
ess, the mixed wastes are blended with coke breeze reductant and binder, pelletized, and furnace 
smelted to recover the contained metals. Although all-pellet heats can be smelted, the recommended 
procedure is to charge the pellets to the production electric arc furnace to replace up to 20 pet 
of the scrap charge. A number of industrial-size (19-st) commercial heats have been successfully 
made with pelletized wastes representing 14 to 19 pet of the furnace charge. Cost evaluation indi- 
cates the process is economically attractive. 



INTRODUCTION 



For technical and economic reasons, the scrap market has tradi- 
tionally concentrated on metallic materials, with the low-grade dusts, 
fumes, solutions, and sludges being normally destined for landfills 
or other available means of disposal. Until recent years, producers 
of such wastes have had little incentive to treat them for metals 
recovery. 

In the domestic specialty steelmaking industry, for instance, 
it is estimated that over 20 million lb Cr and 8 million lb Ni, plus 
other valuable metals, are contained in flue dusts, mill scale, and 
grinding swarfs in a typical year. 

Traditional waste handling and disposition has undergone a sig- 
nificant change over the past decade as a result of the passage in 
1976 of the Resource Conservation and Recovery Act (RCRA) and 
related legislation, which is concerned with the "cradle- to-grave" 
generation, transportation, and disposal of hazardous waste. 
Chromium-bearing electric arc furnace dust has been categorized 
by the U.S. Environmental Protection Agency (EPA) as a hazard- 
ous waste. Other particulates are deemed hazardous if they fail the 
EPA's Extraction Procedure (EP) toxicity test (1) .^ Companies that 
generate hazardous wastes are faced with much-increased costs of 
complying with regulatory standards for storing, transporting, and 
disposing of the wastes. Even costly landfilling may not remain 
an option as regulations banning land disposal of hazardous wastes 
and directives dictating utilization of best available treatment tech- 
nology are being considered seriously by EPA (2). 



'Supervisory metallurgist (research supervisor). 
^Metallurgist. 

Rolla Research Center, Bureau of Mines, Rolla, MO. 

'Italic numbers in parentheses refer to items in the list of references at the end of 
this paper. 



One alternative to disposal of such wastes is to develop 
improved recovery and recycling technology, which can augment 
domestic supplies of critical metals such as Cr and Ni, as well as 
associated metals such as Fe and Mo. 

The Bureau of Mines has conducted research on steelmaking 
and other wastes over a number of years. Much of the research 
(3-6) involved various schemes to remove Zn and Pb from carbon 
steelmaking furnace dusts, using combinations of techniques such 
as sulfation, pelletization, reduction roasting, and specialized fur- 
nacing. 

Bureau recycling research (7) was also conducted on stainless 
steelmaking baghouse dusts, mill scale, and grinding swarf. Phys- 
ical separation and leaching tests proved impractical for segregat- 
ing the metal values. The constituents were too intimately mixed 
for beneficiation separations and too refractory and complex for 
selective leaching. Acid-soluble constituents resulted in a high acid 
consumption. The only method showing promise involved pelletiz- 
ing the wastes with coke breeze reductant and smelting to produce 
a master alloy. Encouraging laboratory results in this earlier research 
led to the making of a 1-st trial heat in an industrial plant (8). 

These earlier heats involved all-pellet charges and production 
of master alloy ingot. Overall metal recoveries were deemed promis- 
ing, with Cr recoveries being consistently somewhat lower than 
the recoveries of Ni and Fe. A preliminary economic evaluation 
was favorable. Further research involved initial tests of adding the 
pelletized wastes to the electric furnace in lieu of part of the con- 
ventional scrap charge (9-1 J). 

As an alternative to in-plant recycling, a master alloy can be 
produced by a centralized waste processing facility. An example 
of this type of facility is the Inmetco plant of the International Nickel 
Co. (INCO), which has in recent years been processing part of the 



51 



stainless steelmaking dusts, mill scale, and grinding swarf being 
generated (12). The treated and smelted compositions are adjusted 
with higher grade materials to produce recyclable Fe-Cr-Ni ingot. 

Fosnacht (13) has compiled a comprehensive bibliography of the 
state-of-the-art of the processing of particulate steelmaking wastes. 

The results reported herein represent efforts of Bureau research 
during the past decade to find technically and economically feasi- 
ble recycling technology for specialty steelmaking wastes— to per- 
mit return of otherwise lost metal values to the steelmaking circuit. 



An in-plant reduction technique has been developed that results in 
the consistent recovery of about 90 pet of the Cr and Mo and well 
over 90 pet of the Ni and Fe from particulate stainless steel wastes. 
Description is given of the work, which began with small labora- 
tory arc furnace melts, as it progressed through a number of approx- 
imately 19-st demonstration heats at Joslyn Stainless Steels, Fort 
Wayne, IN, under a Bureau-industry cooperative effort. Discus- 
sion of a cost evaluation is included. 



PROCEDURE AND RESULTS 



RAW MATERIALS 



Pelletizing 



Waste products evaluated in the research included electric fur- 
nace (EF) baghouse dust, AOD vessel baghouse dust, oily grind- 
ing swarf, and mill scale. The majority of the material was supplied 
by Joslyn Stainless Steels; partial representative analyses are shown 
in table 1 . Samples of baghouse dust and mill scale also were 
obtained from several other companies. Variations noted in com- 
position of the wastes reflect variations in the feed to the furnaces, 
operating procedures, waste collection, and product mix. 

Table 1 .—Partial chemical analyses of typical stainless 
steelmaking waste products, percent 



Waste material 



Fe 



Cr 



Ni 



Mo IVIn Pb 



Zn 



EF dust 27.8 

AOD vessel dust 40.5 

Grinding swarf 61.6 

Mill scale 54.8 



9.3 


2.2 


1.1 


3.6 


0.8 


4.9 


1.1 


3.8 


.7 


5.5 


.6 


.8 


1.7 


6.8 


1.2 


1.0 


.1 


<.1 


8.6 


3.9 


.5 


.8 


.1 


<.1 



Arc furnace and AOD vessel baghouse dusts are typically about 
90 pet minus 400 mesh. They consist essentially of complex oxides 
involving metals such as Fe, Cr, Ni, Mo, Mn, and Zn, with a num- 
ber of other constituents involving elements such as Ca, Mg, K, 
Na, Si, S, C, etc. Mill scale, which is produced during blooming 
of heated billets, also is composed principally of metal oxides. Mill 
scale is mostly powder, but it is typically much coarser than bag- 
house dusts. Some larger chunks of mill scale commonly are present. 

Swarfs are generated by the surface grinding of billets, slabs, 
and bars. Some of the dry swarf resulting from bUlet grinding, which 
is less oxidized than baghouse dusts and mill scale, is recycled 
directly back to the electric furnace. The oily swarf used in these 
experiments was specifically from the centerless grinding of vari- 
ous sized rods. It consisted of many small, partially oxidized frag- 
ments containing entrained oxide and carbide particles from the 
grinding wheels. It was often still wet from cutting oils used as a 
coolant during the grinding. The particle size of swarf is typically 
somewhat coarser than the finer fractions of mill scale. 

The coke breeze obtained from a steel plant ranged downward 
in size from about U in; it was about 85 pet C. The portland cement 
used as a binder typically has a very fine particle size. 

LABORATORY EXPERIMENTS 

Laboratory experiments were conducted in which various waste 
mixtures pelletized with coke breeze provided approximately 100-lb 
charges for smelting in an arc furnace. 



Pelletizing was selected as the most practical and economic 
means for agglomeration of the wastes. The four steelmaking wastes 
(table 1), in proportions consistent with their rates of generation, 
were mixed with coke breeze reductant and cement binder and 
blended into a composite mixture. For a plant that is collecting all 
four wastes, this composition might typically represent some 15 
pet of AOD dust, 15 to 25 pet each of EF dust and mill scale, and 
about 40 pet grinding swarf. Substantial variation can, of course, 
be expected. Coke breeze, added as about 10 wt pet to this waste 
blend, was selected as a logical carbon reductant because it is essen- 
tially in a powdered condition and is an accessible carbon source 
in the steelmaking industry. With a wide particle size range for steel- 
making wastes, a binder may be necessary, particularly if propor- 
tions of coarser sizes are relatively high. An addition of about 4 
pet cement was found to be an effective binder for the waste mixes 
evaluated. The fine particle size of the cement aids pellet forma- 
tion prior to the hardening reaction. 

When raw, somewhat caked coke breeze is used, it must be 
dried and crushed to pass an intermediate size screen, such as 35 
mesh, before adding it to the mixture. The mill scale was also dried, 
and then passed through crushing rolls until >70 pet passed a 
35-mesh screen— usually two passes through the rolls. Only the 
minus 35-mesh fraction was blended into the pellet mix used in 
smelting tests; the oversize fraction was saved and added with the 
pellets as furnace charge. A flow diagram of this procedure is 
presented in figure 1. 

The four-waste mixture was pelletized in a 36-in-diam drum 
pelletizer. The blended mix required the addition of about 12 pet 
water in order to form %- to %-in-diam pellets. The pellets produced 
were first air dried for 24 h and then oven dried at 250° F for 6 h. 
Drying at higher temperatures tended to result in spalling. The 
pellets had 5- to 30-lb crushing strength, which was sufficient for 
the limited amount of handling required. When the wastes listed 
in table 1 were mixed in the indicated proportion and pelletized, 
the resultant pellets plus oversize mill scale analyzed roughly, in 
percent, 10 Cr, 4 Ni, 1 Mo, and 2 Mn. 

Experience showed, however, that it is not always necessary 
to oven dry the pelletized waste mix. Most of the laboratory heats 
made in the latter part of the testing program were made using pellets 
that were air dried only. The experiments included adding pellets 
to the furnace that were only 2 h old and very wet. These were 
rolled at a uniform rate down a conveyor and dropped through the 
furnace top onto the melt surface. As much as 13 pet of the melt 
charge (the most tried) was added in this manner without problem. 

Other efforts were made to simplify the pelletizing procedure. 
For some trials, the cement binder was completely eliminated from 
the pellet mix (could decrease slag volume). No difficulty was 



52 



Col<e breeze 

(minus 35-mesh) 

10 pet 



Cement 
4 pet 



Electric 

furnace 

dust 

17 pet 



AOD 

vessel dust 
12 pet 



Grinding 
swarf 
40pct 



Mill scale 
17 pet 



n r 



Blender 



Water- 



-Minus 35 meslr 



Screen 
3/4 -in 
35 - mesh 



Plus 35 mesfi 



Pelletizer 



Air dry 24 h, 
oven dry 6 h at 250°F 



Rol I crusher 
2 posses 



Plus 3/4 in 



Screen 
35-mesh 



Plus 35 nnesh 



Finished pellets 



Mil 



sea le 



Figure 1 .—Flow diagram for laboratory agglomeration of four stainless steelmaking wastes by pelletizing with coke breeze reduc- 
tant and cement binder. Pellets of somewhat reduced but generally adequate strength can also be produced by air drying only. 



encountered in making pellets; however, after storage for several 
weeks, the pellets tended to spall and powder. For extended stor- 
age, some binder appears necessary. Adding 1 to 2 pet bentonite 
resulted in acceptable pellets that were stable for longer periods 
than when no cement was added, although not as stable as when 
the binder was cement. It was also found that the mill scale could 
be crushed and screened through 20- or even 10-mesh (rather than 
35-mesh) screens with no apparent effect on pellet properties. 

Laboratory Smelting of Stainless 
Steelmaking Wastes 

Earlier tests in an induction furnace (7) indicated that numer- 
ous combinations of pellet compositions could be smelted, with the 
carbon in the coke breeze reducing essentially all the oxides of Fe 
and Ni and a considerable part of the Cr oxide, roughly about one- 
half. The Mn oxide present was also readily reduced. It was 
observed that ferrosilicon could then be added to the molten bath 
to scavenge much of the remaining Cr from the slag. 

The procedure for 100-lb arc furnace melts began with pre- 
heating the furnace refractories for 1 h, followed by the charging 
of 90 to 95 lb of pellets over a 45-min period. The 5- to 10-lb por- 
tion of loose (plus 35-mesh) mill scale was then added, and the fur- 
nace temperature was brought to at least 2,950° F. At this point, 
3.0 pet Si, as ferrosilicon (75 pet Si), was added to the meh, which 
was stirred vigorously to enhance mixing and effect additional reduc- 
tion of Cr oxide from the slag. The melt was heated to a tempera- 
ture of 2,950° to 3,100° F before tapping (fig. 2). The ingots 
accounted for about 60 pet metal yield from the wastes charged. 
The slag represented some 10 to 15 pet of the original charge weight. 



Further experiments indicated that 0.3 pet Al (as percentage 
of charge weight), added as shot, accomplished roughly the same 
amount of Cr and Ni reduction as the 3.0-pct Si addition. Table 
2 shows results of the tests comparing the use of Fe-Si, Fe-Si-Al, 
and Al. The increased reduction associated with the use of Al shot 
may be due to the lower melting point or the higher thermodynamic 
activity of pure Al; the same nominal Al addition as the Al-bearing 
ferrosilicon was less effective, even in combination with the con- 
tained Si. For the 3-pct Si addition, most of the Si reported to the 
ingot. 

Table 2. — Recovery of chromium and nickel from pelletized 
wastes after the smelting addition of various reductants^ 



Reductant added, 


Ingot content, pet 


Recovery, pet 


pet 


Cr Ni Si 


Cr Ni 


3 2Si 

0.6 Si, 0.3 3AI 

0.3 "Al 


14.2 71. 6.9 

. ... 15.1 6.9 2.3 
15.5 7.1 .7 


91.5 99.3 
84.7 90.5 
93.4 97.8 



1 In addition to 10 pet eoke breeze in pellets. 

2 As ferrosilieon (75 pet Si). 

3 As aluminum ferrosilicon (20 AI-40 Fe-40 Si). 
"As aluminum shot. 



Other heats were made without Si or Al reductants. These melts 
were held similarly for 20 to 30 min before tapping, within the range 
2,850° to 3,050° F. Only those held 30 min at 3,050° F showed 
improved Cr recovery relative to that for the lower temperatures. 
The reduction of the Cr oxide is related to time and temperature, 
as well as the type and amount of reductant. 



53 





■i 









.y' 



Figure 2.— Laboratory arc furnace being tapped after smelting a charge of pelletized stainless steelmaking wastes. 



54 



Some stainless steel producers had indicated that they had been 
stockpiling mill scale. Samples were obtained from five compa- 
nies. Partial analyses are shown in table 3. The scale from com- 
pany C originated from ferritic stainless steel production; the others 
were from austenitic stainless steel production. The mill scales were 
mixed with other wastes and pelletized in two approximate com- 
positions, as shown in table 4. 

Although arbitrary, the proportions for the low-scale compo- 
sition are intended to represent an approximate generation rate in 
a plant producing and segregating all four wastes. The high-scale 
composition reflects a situation whereby stockpiled mill scale would 
be used up at a faster-than-generated rate. Small-scale pelletizing 
tests indicated that pellets containing as much as 55 pet mill scale 
could be made, but 35 pet was a more practical value from the stand- 
point of a higher proportion of fines present and better pellet for- 
mation and strength. 

The mixtures were prepared, blended to the compositions 
shown for low and high mill scale, and arc furnace melted as 
described previously. The slag was reduced with ferrosilicon. The 
results obtained for the group representing low mill scale (17.4 pet) 
pellets are shown in table 5. Manganese recoveries were 85 pet 
or more. 

The smelting experiments with the low mill scale pellets indi- 
cated that 82 to 92 pet of the Cr could be readily recovered. Addi- 
tional reductant and/or additional heating would increase the 
recoveries. Nickel recovery was, as expected, higher than the Cr 
recovery. The results of the smelting trials on high mill scale 

Table 3.— Partial analyses of stainless 
steel mill scale samples, percent 



Company 
designation 



Fe 



Cr Ni Mo 



A 54.8 8.6 3.9 0.48 

B 51 .9 11 .2 7.0 .47 

C 62.7 7.6 .5 .22 

D 49.4 8.4 3.7 .16 

E 56.2 9.1 3.0 .66 



Table 4.— Composition of pellets with low 
and high mill scale content, percent 



Connponent 



Low mill 
scale 



High mill 
scale 



Mill scale 

EF dust 

AOD dust 

Grinding swarf 

Coke breeze 

Cement 

Total 100.0 



17.4 


35.0 


17.4 


12.6 


13.0 


9.5 


39.1 


28.0 


8.7 


10.5 


4.4 


4.4 



100.0 



(35.0 pet) mixtures indicated a Cr recovery 5 to 10 pet lower than 
for low (17.4 pet) mill scale mixtures. This was believed due, in 
part, to the fact that the higher mill scale content resulted in a some- 
what increased slag volume. That is, even with the same Cr slag 
solubility per unit volume, more total Cr partitioned to the higher 
volume slag. 

When other factors are held constant, the Cr solubility in slag 
decreases with increased slag basicity (J4), with slag basicity being 
defined as the ratio of percentage CaO plus MgO to SiOj. The 
increased slag volume and weight brought about by the increased 
lime addition needed to increase the basicity may, however, offset 
any gain in reduced Cr solubility when considering the total Cr in 
the slag. 

It should be noted that the laboratory smelting tests involved 
a charge consisting of 90 to 95 pet pellets and 5 to 10 pet loose 
mill scale. Some producers might have furnace capacity available 
to intermittently melt all-pellet charges for metering of hot metal 
to production heats. This method could also help reduce large back- 
logs of wastes. However, in normal industrial practice, it is antici- 
pated that only 10 to 20 pet of the furnace charge would consist 
of pelletized wastes, to replace a portion of the scrap charge. This 
is the most economical and least energy-intensive approach and can 
readily match or exceed waste generation rates. Excess reductant 
normally available in production heats can also assist metal recov- 
eries from the wastes. Under these circumstances, the slag genera- 
tion would not be as much as indicated for the heats in table 5 but 
should be near to, although perhaps toward the high end of, nor- 
mal ranges. 

It remains to be conclusively demonstrated to what extent some 
of the minor metals such as Zn and Pb in the EF dust will build 
up in the baghouse dust with repeated pelletizing and refumacing. 
The longer term concentration of minor elements (rate and extent) 
in recycled furnace dust will not be clarified until results are rejxjrted 
for campaigns extending over a substantial number of heats, with 
carefiil analysis of all charge and product materials. 

There have been some extended campaigns to recycle baghouse 
dusts to the furnace without added carbon or other reductant. For 
these, the main objective has been to decrease the volume of haz- 
ardous waste requiring disposition, rather than gaining metal recov- 
ery. One firm has been recycling pelletized minimill carbon 
steelmaking furnace dust in this manner (15). Even without reduc- 
tant added with the pellets, some increased Fe yield has apparently 
been realized in these instances, evidently by supplying some 
"equilibrium" slag iron oxide requirements from the furnace dust 
rather than by Fe oxidation from the melt. Some limited reduction 
may be expected from excess reductant (Si and C) present in the 
melt. Furnace dusts generated from heats to which pelletized fur- 
nace dusts have been recycled are characteristically enriched in Zn 
and other metals that are volatile or form volatile species at steel- 
making temperatures. For further information on the composition 
and nature of carbon and stainless steelmaking furnace dusts, the 
interested reader is directed to the final report resulting from the 



Table 5.— Results of smelting pelletized waste mixtures containing mill 
scale^ obtained from five stainless steel producers 



Company 
designation 



Ingot analysis, pet 



Recovery, pet 



Cr 



Ni 



Mo 



Cr 



Ni 



Mo 



Ingot as Slag as 

percent percent 

of cfiarge of cfiarge 



A 16.0 

B 16.1 

14.5 

D 14.9 

E 14.9 



7.2 


1.00 


92.6 


100.0 


74.5 


60.0 


14.0 


7.7 


.80 


87.7 


90.7 


91.9 


47.9 


13.0 


5.9 


.79 


86.1 


97.9 


91.5 


56.9 


16.9 


7.1 


.83 


82.3 


93.2 


63.6 


55.9 


19.3 


6.2 


.92 


93.0 


99.0 


100.0 


63.4 


19.0 



M7.4 pet mill scale in thie pellets; other ingredients were (in percent) 17.4 EF dust, 13.0 AOD dust, 39.1 grinding swarf, 8.7 
coke breeze, and 4.4 cement. 



55 



comprehensive arc furnace dust program conducted for the U.S. 
Department of Commerce by Lehigh University and the Bureau 
of Mines, with AISI collaboration and coordination (16). 

LARGE-SCALE DEMONSTRATION HEATS 
All-Pellet Heats 

The procedures derived in the laboratory experiments were tried 
on a larger scale in the form of a 1-st heat of the pelletized waste 
mixture (8), which was made by Union Carbide Corp., Niagara 
Falls, NY. When it was evident that recoveries from this heat were 
sufficiently good, plans were formulated for substantially larger 
scale industrial trials. 

The first of these large-scale demonstration trials was an all- 
pellet heat of approximately 12.5 st made in the production plant 
of Joslyn Stainless Steels. The wastes, generated during normal 
operations at Joslyn, were pelletized by a contractor in accordance 
with specifications developed at the Bureau's Rolla (MO) Research 
Center. The pellets ranged from %- to 1-in diam and represented 
the mixture indicated earlier as low (17.4 pet) mill scale. The pellets 
analyzed, in percent, 42.5 Fe, 9.6 Cr, 4.0 Ni, and 0.7 Mo. 

The charge to the furnace consisted of approximately 20,900 
lb of pellets and 1,100 lb of oversize mill scale along with 0.5 st 
of lime. Several hundred pounds of steel punchings had been added 
at the outset to help strike an arc. When the charge was fully mol- 
ten, the melt was sampled over a 30-min period. An addition of 
1,320 lb of ferrosilicon (50 pet Si) was made to reduce chromium 
oxide remaining in the slag. Improved contact between the ferrosili- 
con and slag was achieved by stirring with an argon lance. The 
melt was tapped through the slag into a ladle from which it was 
poured into 14- by 14-in molds. The master alloy ingot produced 
weighed 12,300 lb and analyzed, in percent, 76.7 Fe, 11.8 Cr, 6.5 
Ni, 0.8 Mo, 0.9 Mn, 4.3 Si, and 3.2 C. These values represented 
recoveries of 86. 1, 68.7, and 92.0 pet of the Fe, Cr, and Ni, respec- 
tively. As shown in table 6, power consumption was 1 .05 kWh/lb 
of metal tapped. 

The Cr recovery was lower than that experienced in smaller 
scale tests but was considered good for a single-heat experiment. 
Temperature was probably responsible for the lower recovery. After 

Table 6.— Confiparison of charge weight, nfietal tapped, 

and energy requirements for 

Joslyn demonstration heats 1 and 2 





Charge, lb 


Waste, 
pet 


Metal 
tapped, 

lb 


Power con- 


Heat Type of heat 


Total Waste 


sumption,' 
kWh 


1 All pellet 

2 Scrap plus heat 1 

master alloy. . 


25,385 22,010 
42,025 7,980 


86.7 
19.0 


12,320 
39,700 


1.05 
.254 



the ferrosilicon addition, the bath temperature of 2,750° F was some 
200° F below the temperature of most of the laboratory heats. 
Nevertheless, Ni, the metal of highest total value, reported to the 
ingot as expected. 

An 8,0()0-lb portion of the master alloy ingot from the all-pellet 
heat was incorporated into a commercial 19-st heat (heat 2, table 
6) of AISI Type 316 stainless steel. The master alloy was com- 
pletely compatible with the balance of the charge, principally stain- 
less scrap. The heat was completed within the expected time, power 
consumed was normal, and there were no problems in either the 
arc furnace melting or AOD refining. The remainder of the ingot 
material was used in another commercial heat with equally good 
results. 

Pellet-Plus-Scrap Heats 

At this point, rather than optimizing the making of all-pellet 
heats, it was decided to go directly to the overall simpler and more 
economical introduction of waste pellets into the production arc ftir- 
nace in lieu of part of the normal scrap charge. It was judged that 
the advantages of this approach would be ( 1 ) lower energy con- 
sumption, (2) simpler processing without producing intermediate 
master alloy ingots requiring remelting, and (3) variation of the 
quantity and composition of waste pellets as dictated by needs. It 
was realized that careful attention would have to be given to the 
furnace charge to accommodate the pellet waste ingredients and 
their products. However, since the makeup of furnace charges is 
commonly computer calculated, programs can be readily adjusted 
to account for this unconventional raw material. 

Five Type 316 stainless steel heats were made in a 19-st produc- 
tion furnace, in which pelletized wastes (high or low scale) con- 
stituted 14 to 19 pet of the nominal charge. The remaining pellets 
from the Joslyn all-pellet heat, hereafter referred to as low-scale 
pellets, were used for two of these heats. Tonnage quantities of 
pellets also were produced by a contractor with substantially aug- 
mented percentage of mill scale. This high-scale composition added 
in the last three heats was tested to simulate consumption of sub- 
stantial quantities of stockpiled mill scale. The makeup of the high- 
scale pellets was as follows, in percent: mill scale, 30; EF dust, 
13; AOD dust, 9; grinding swarf, 33; coke breeze, 11; and cement, 
4. The pellet composition (including oversize mill scale) was, in 
percent, 39.2 Fe, 8.9 Cr, 3.7 Ni, and 0.5 Mo. 

The pellets were added to the arc furnace concurrently with 
the stainless scrap, making it unnecessary to backcharge. For all 
heats, only that quantity of ferrosilicon normally added to "quiet" 
such scrap melts was added. This ranged from to 300 lb. The 
slag volume in each case was considered within the normal range 
for production heats. (With an extended campaign, some increased 
slag volume can be expected when partly replacing relatively clean 
scrap with pelletized wastes — other factors equal.) The pertinent 
statistics for the five heats are presented in table 7. All of the heats 



Table 7. — Comparison of charge weight, metal tapped, and energy 

requirements for Joslyn demonstration heats 3 through 7 

(pellets partly replace scrap) 





Type of heat 


Charge, 


lb 


Waste, 


Metal 
tapped. 


Power con- 
sumption,' 


Tap 


Heat 






temp. 






Total 


Waste 


pet 


lb 


kWh 


° F 


3.. 


Scrap plus low- 
scale pellets. 


41,790 


5,970 


14.3 


38,500 


0.244 


2,920 


4. . 


. . .do 


41 ,600 


7,870 


18.9 


36,500 


.260 


2,960 


5.. 


Scrap plus high- 
scale pellets. 


41 ,940 


6,300 


15.0 


38,000 


.263 


3,000 


6.. 


. . .do 


43,490 


6,250 


14.4 


36,600 


.259 


2,950 


7. . 


. . .do 


42,900 


8,200 


19.0 


36,300 


.261 


2,940 


' Per pound of metal tapped. 















56 



met required specifications after processing through the AOD ves- 
sel and were marketed as commercial bar, rod, or forging ingot. 
Table 8 gives the recoveries of Cr, Ni, and Mo, which, on the aver- 
age, were considered equivalent to the customary values for all- 
scrap heats. Iron recoveries were consistently substantially > 90 pet. 

Metal Value of Pelletized Wastes 

Technically and mechanically, this recycling scheme has been 
shown to work well. The question naturally arises as to whether 
it is also economical. The Bureau's Process Evaluation Group con- 
ducted internal studies of capital and operating costs for a plant addi- 
tion producing pellets from wastes such as flue dusts, mill scale, 
and/or oily swarf. The estimated fixed capital cost for a 15-st/d 
pelletizing capacity was approximately $974,000 for oven drying 
of pellets and approximately $560,000 for air drying. Estimated 
annual operating costs per ton of pellets based on one-shift-per- 
day, 5-day -per-week operation (20-yr life) was approximately 
$117/st for oven drying and $40/st for air drying. 

This can be contrasted to the contained value of the Cr, Ni, 
Mo, and Mn in the pellets. In addition to the particular propor- 
tions of the metals in the wastes, the value depends on the current 
price of ferroalloys or of appropriate scrap for which the pellets 



Table 8.— Recovery of Cr, Ni, and Mo from commercial 
pellet-plus-scrap heats 3 through 7, percent 



Heat 


Type of heat 


Cr 


Ni 


Mo 


3 . . 

4 . . 

5 .. 

6 .. 
7.. 


Scrap plus 14 pet low-scale pellets. . . . 
Scrap plus 19 pet low-scale pellets. . . . 
Scrap plus 15 pet high-scale pellets. . . 
Scrap plus 14 pet high-scale pellets. . . 
Scrap plus 19 pet high-scale pellets. . . 


93.0 
93.7 
97.3 
90.8 
91.4 


99.7 
89.7 
99.1 
91.9 
92.3 


99.9 
95.0 
93.2 
82.8 
84.5 



may substitute, particularly stainless steel (18 Cr-8 Ni) scrap. In 
some periods, a charge is also added for Fe units; this has not been 
the case when the scrap market is relatively depressed. On the basis 
of the Cr, Ni, Mo, and Mn for a waste mixture similar to those 
described for low-scale pellets (in percent, 9.5 Cr, 4.0 Ni, 0.8 Mo, 
and 2.0 Mn), with approximate contained metal values of $0.42/lb 
for Cr, $2.10/lb for Ni, $3.20/lb for Mo, and $0.33/lb for Mn 
(mid- 1987 ferroalloy contained-metal values), the pellets would have 
a value of over $3I0/st. Deducting the approximate net operating 
cost of $1 17/st for oven drying or $40/st for air drying indicates 
a net gain of some $0.10 or $0.14/lb, respectively— a significant 
economical potential. 



CONCLUSIONS 



It has been shown that stainless steelmaking wastes such as flue 
dusts, mill scale, and grinding swarf can be pelletized and reduced 
in the arc furnace. This provides a means of recovering the con- 
tained scarce and valuable metals, while coincidentally solving prob- 
lems of storage and waste disposal. The recovery procedure utilizes 
the reduction of metal oxides with carbon during the arc furnace 
melting, followed by a scavenging slag reduction of additional chro- 
mium oxide with ferrosilicon or Al. 

One variation of the processing involves preparation of all-pellet 
smelting heats to produce master alloy ingot for recycle. This may 
be appropriate if furnace capacity is available in slack periods and 
large waste backlogs exist. 

Alternatively, and recommended as being more economical, 
the waste-bearing pellets can be added directly to the arc furnace 
as some 10 to 20 pet of the total charge for production heats in 
lieu of part of the usual scrap or alloy charge required. The addi- 
tion rate will depend on factors such as the rate of waste genera- 
tion, waste backlog accumulation, and alloy product mix at a 
particular plant. 

The dusts, scale, and swarf can be mixed and pelletized with 
little difficulty, providing both a means for adding carbon to the 
mix and a vehicle for charging to the furnace. Numerous varied 
combinations of pellet mix have been shown to be possible. Only 
conventional equipment is needed for agglomeration. 



Usual recoveries of substantially greater than 90 pet of the Ni 
and Fe have been attained, and some 90 pet of the Cr and Mo appear 
consistently recoverable with proper control of variables. Other 
metals such as Mn are coincidentally recovered. Conventional arc 
furnaces were used throughout the testing. 

The fact that this technology is readily transferable to an indus- 
trial scale was shown by the successful making of a number of 
demonstration heats ranging in size from 12.5 st for an all-pellet 
heat to about 19 st for a series of pellet-plus-scrap heats. The mas- 
ter alloy ingot from the all-pellet demonstration heat was used to 
make commercial stainless steel products without difficulty. No 
problems were encountered in these commercial stainless produc- 
tion heats to which up to 19 pet pellets were added in lieu of the 
normal scrap charge. 

When the particular waste combination outlined in this paper 
is pelletized for recycle as a scrap substitute charge material, the 
pellets have a net value of more than $O.IO/lb for oven drying or 
$0.14/lb for air drying, which implies an economically viable 
process. 

Test results also indicate that a wide compositional variation 
of specialty steelmaking wastes can be incorporated into pellets for 
furnace charging. 



57 



REFERENCES 



1. Federal Register. U.S. Environmental Protection Agency. Hazard- 
ous Waste Regulations. V. 45, No. 98, 1980, p. 33127. 

2. Price, L. E. Tensions Mount in the EAF Dust Bowl. 33 Met. Produc- 
ing, Feb. 1986, pp. 38-41. 

3. Barnard, P. G., A. G. Starliper, W. M. Dressel, and M. M. Fine. 
Recycling of Steelmaking Dusts. BuMines TPR 52, 1972, 10 pp. 

4. Dressel, W. M., P. G. Barnard, and M. M. Fine. Removal of Lead 
and Zinc and the Production of Prereduced Pellets From Iron and Steel- 
making Wastes. BuMines RI 7927, 1974, 15 pp. 

5. Higley, L. W., Jr., and M. M. Fine. Electric Furnace Steelmaking 
Dusts— A Zinc Raw Material. BuMines RI 8209, 1977, 15 pp. 

6. Higley, L. W., Jr., and H. H. Fukubayashi. Method for Recovery 
of Zinc and Lead From Electric Furnace Steelmaking Dusts. Paper in 
Proceedings of the Fourth Mineral Waste Utilization Symposium. IIT Res. 
Inst., Chicago, IL, 1974, pp. 295-302. 

7. Powell, H. E., W. M. Dressel, and R. L. Crosby. Converting Stain- 
less Steel Furnace Flue Dusts and Wastes to a Recyclable Alloy. BuMines 
RI 8039, 1975, 24 pp. 

8. Barnard, P. G., W. M. Dressel, and M. M. Fine. Arc Furnace Recy- 
cling of Chromium-Nickel From Stainless Steel Wastes. BuMines RI 8218, 
1977, 10 pp. 

9. Higley, L. W., Jr., L. A. Neumeier, M. M. Fine, and J. C. Hart- 
man. Stainless Steel Waste Recovery System Perfected by Bureau of Mines 
Research. 33 Met. Producing, Nov. 1979, pp. 57-59. 



10. Higley, L. W., Jr., R. L. Crosby, and L. A. Neumeier. In-Plant 
Recycling of Stainless and Other Specialty Steelmaking Wastes. BuMines 
RI 8724, 1982, 16 pp. 

11. Neumeier, L. A., and M. J. Adam. In-Plant Recycling of Chromium- 
Bearing Specialty Steelmaking Wastes. Paper in Chromium-Chromite: 
Bureau of Mines Assessment and Research. Proceedings of Bureau of Mines 
Briefing Held at Oregon State University, Corvallis, OR, June 4-5, 1985, 
BuMines IC 9087, 1986, pp. 85-91. 

12. Pargeter, J. K. Operating Experience With the Inmetco Process for 
the Recovery of Stainless Steelmaking Wastes. Paper in Proceedings of the 
Seventh Mineral Waste Utilization Symposium. IIT Res. Inst., Chicago, 
IL, 1980, pp. 118-126. 

13. Fosnacht, D. R. Recycling of Ferrous Steel Plant Fines: State-of- 
the-Art. Iron and Steelmaker, v. 8, No. 4, Apr. 1981, pp. 22-26. 

14. Peckner, D., and I. M. Bernstein. Handbook of Stainless Steels. 
McGraw-Hill, 1977, pp. 3-1-3-35. 

15. Mueller, C. P. Recovery of Metallics From Specialty Steel Slags 
and Wastes. Pres. at AISI Symposium on Recovery of Alloys From Spe- 
cialty Steel Wastes, Pittsburgh, PA, Oct. 21-22, 1981; information availa- 
ble from International Mill Service, Inc., Philadelphia, PA. 

16. Lehigh University. Characterization, Recovery and Recycling of Elec- 
tric Arc Furnace Dusts. Final rep. prepared for U.S. Dep. Commerce under 
project 99-26-09885-10, Feb. 1982, 313 pp.; NTIS PB 82-182585. 



58 



USING WASTES AS A SOURCE OF ZINC FOR ELECTROGALVANIZING 



By V. R. Milleri 



ABSTRACT 

The Bureau of Mines investigated the use of Zn extracted from electric arc furnace (EAF) 
dust as a source of Zn for electrogalvanizing. The prepared sulfate electrolytes were used to coat 
steel sheet in flow cells at current densities up to 150 A/dm^ to provide 90-g/m2 Zn deposits. Elec- 
trochemical and physical properties of waste-derived coatings were compared with those of coat- 
ings produced from electrolytes prepared from pure ZnO and from an industrial process. These 
studies showed the properties to be similar in most cases. 



INTRODUCTION 



A dramatic increase in the demand for electrogalvanized steel 
has resulted from the construction, appliance, and especially the 
automotive industry needs (1).^ The amount of slab zinc used in 
1986 for galvanized steel was estimated to be 47 pet of the total 
slab zinc consumption of 975,000 mt. An estimated 73 pet or 
710,000 mt of this slab zinc was imported. Over 20 yr ago, it was 
estimated that 235,000 st of zinc could be recovered from domes- 
tic stack dusts (2). In 1980, 75,000 st of Zn was contained in EAF 
dust alone (3). This quantity of Zn was contained in about 400,000 
st of EAF dust which has been declared a hazardous waste by the 
Environmental Protection Agency (4). 

Several studies and surveys have been carried out in recent years 
to examine the options for recovery of resources from EAF dust. 
Several processes, both pyrometallurgical and hydrometallurgical, 
have been proposed for treating EAF dust and these have been well 
documented in the literature. The bulk of these processes were aimed 
at recovering pure metals such as Zn. In the case of Zn, this requires 
stringent purification to produce electrolytes with impurities in the 



parts per billion range. However, for electrogalvanizing, some of 
the impurities have been shown to be beneficial (5). Cobalt and 
chromium at low concentrations in electrogalvanizing baths have 
produced more corrosion-resistant coatings and Cd has enhanced 
wire drawability. 

The Bureau of Mines demonstrated in previous research the 
feasibility of using Zn recovered from brass smelter flue dust to 
electrogalvanize steel wire in an industrial pUot plant (6). This paper 
describes research on electrogalvanizing steel sheet with electro- 
lytes produced from EAF dust and an EAF dust oxide fume 
produced from flash smelting the dust. The oxide fume differs from 
the EAF dust in that it contains the more volatile metals. Solutions 
from leaching the two products were purified as needed for elec- 
trogalvanizing and used in flow cells designed to simulate the 
hydrodynamics of an industrial line. Properties of deposits made 
with waste dust electrolytes were compared with the properties of 
deposits made using pure ZnS04 as well as industrial deposits. 



EXPERIMENTAL 



WASTE MATERIAL AND ELECTROLYTE 

Table 1 contains a partial chemical analysis of the dusts used 
in the experimental work. The electrolyte from the EAF dust was 
prepared by mixing concentrated H2SO4 with the dust to sulfate 
it, then water leaching. The leach step involved mixing the sulfated 
dust with de-ionized water at 90° C for 1 h. The mixture was 
filtered; the liquor was reserved for electrolyte purification and the 
filter cake was combined with pure water for the wash step. The 



Table 1.— Partial chemical analysis of dusts used, weight percent 



' Supervisory research physicist, Rolla Research Center, Bureau of Mines, 
Rolla, MO. 

^ Italic numbers in parentheses refer to items in the list of references at the end 
of this paper. 



Element 



Oxide fume from 
flash smelter 



Zn . 

Fe. 

Pb . 

Cd . 

Cu . 

Mn 

Mg, 

Ca . 

CI. 

F.. 

Cr. 



Electric arc furnace 
(EAF) dust 



35.4 


31.3 


7.5 


19.6 


6.5 


5.2 


.8 


.1 


.6 


.2 


.9 


2.8 


.4 


1.3 


.9 


3.6 


6.8 


3.5 


2.5 


.1 


.2 


.3 



wash step consisted of 30-min agitation at ambient temperature fol- 
lowed by filtration. The filter cake was then set aside for sampling 
and analysis. The wash water was returned to the leach reactors 
for starting the leach cycle on the next batch of sulfated dust. Typical 
combined extractions of Zn and Fe in the two steps were 92 and 
11 pet, respectively. 

The oxide fume electrolyte was prepared by direct H2SO4 leach- 
ing of the fume. The procedure was similar to that for the sulfated 
dust except that 450-g/L H2SO4 solution was used rather than de- 
ionized water. The wash step utilized pure water that was used, 
following filtration, for making the next acid leach solution. Typi- 
cal leach liquor compositions are listed in table 2. The necessity 
of a degree of purification for the solutions is also evident when 
table 2 is examined in light of the need for low levels of Fe, Cu, 
Cd, CI, and F. 

Table 2.— Typical chemical analyses of leach solutions from 

oxide fume from flash smelter and electric arc furnace dust, 

grams per liter 



Element 



Oxide fume 



EAF dust 



Zn . 

Fe. 

Cu 

Cd 

CI. 

F.. 

Pb . 



58 


155 


35.3 


10.1 


3.0 


<.01 


.5 


.4 


11.7 


18.6 


5.2 


.6 


<.01 


<.01 



The Fe concentration was reduced to less than 0.1 g/L using 
a phosphate purification technique. The technique consisted of 
reducing the pH to 1.0, oxidizing the Fe with hydrogen peroxide, 
adding H3PO4 as a phosphate ion source, and raising the pH to 1.9 
with lime. The Fe was subsequently precipitated as readily filtera- 
ble FeP04. The CI levels were reduced using CUSO4 and metallic 
Cu powder at 90° C, pH 2, for 2 h to form CuCl. Copper, cad- 
mium, and lead were removed by cementation with Zn dust. Table 
3 lists the chemical compositions of the electrolytes that were used 
to electrogalvanize the steel sheet. The CI content of the oxide fume 
was high because of insufficient purification, which was not detected 
because of analytic error. It was detected when the electrolyte was 
rechecked after electrogalvanizing. Table 3 includes the pure zinc 
sulfate electrolyte that was prepared by dissolving French process 
ZnO in H2SO4 and H2O. Zinc dust was added to the hot solution 
to insure the removal of Cd and Pb. 

Table 3.— Electrolyte compositions used for electrogalvanizing 
steel sheet, grams per liter 



Element 



Pure ZnSOj EAF dust 



Oxide fume 



Zn . 
Fe. 
Cu . 
Cd . 
Pb . 
CI. 
Co. 
Ni. 
Mn. 



99 


91 


92 


<.001 


.025 


.050 


<.001 


.004 


.004 


<.001 


<.001 


.002 


<.002 


.002 


.002 


<.001 


.035 


1.510 


<.001 


<.001 


.230 


<.001 


<.001 


.140 


<.001 


2.94 


1.41 



59 



vent, rinsed with ethanol, and dried in an airstream. The sheets 
were weighed to the nearest 0.0001 g. The sheets were then elec- 
trocleaned anodically in commercial alkaline cleaner for steel. 
Experimentation showed no significant change in weight, < 0.001 g, 
before and after electrocleaning the sheets. The procedure follow- 
ing electrocleaning was to rinse the sheets in water, dip them in 
a 10-pct H2SO4 solution for 15 s to activate the surface, rinse in 
H2O, place in the flow cell, and start the flow of electrolyte. 



ELECTROGALVANIZING 

Electrogalvanizing was done in two different size flow cells 
that used pumps to move the electrolyte past the stationary elec- 
trodes to simulate a moving line. The small cell used 4- by 5-cm 
sheet to yield a coated area of 20 cm^. The anodes were Pb-1 pet 
Ag and were also 4 by 5 cm in size. The electrode spacing was 
9 mm and with available flow rates, velocities at the cathode sur- 
face of up to 6 m/s were possible. Face velocities were used for 
flow rates to allow comparison with industrial electrogalvanizing 
line speeds. Five liters of electrolyte was required for operation 
of the pump to produce the desired flow rates. 

The large cell coated 10- by 20-cm sheets and required 100 L 
of electrolyte. It also used Pb-1 pet Ag anodes with the electrode 
spacing being 9 mm. Time and current were adjusted for the cur- 
rent densities of 50 and 150 A/dm^ to produce coatings of about 
90 g/m^. After the sheets were electrogalvanized, they were rinsed 
in H2O, then ethanol, dried in air, and weighed to determine the 
amount of coating. 



EXPERIMENTAL DESIGN 

In conducting the research on the pure zinc sulfate electrolyte, 
the variables that were considered were current density, tempera- 
ture, acid concentration (pH), Zn concentration, and face velocity 
of the electrolyte. Table 4 lists the variables and the upper and lower 
limits used in the experiments. A factorial design was used to gain 
the maximum information with a minimum of experimental work. 
The responses analyzed were current efficiency and preferred crys- 
tallographic orientation. Selected specimens were also subjected to 
electrochemical corrosion evaluation in addition to formability 
testing. 



Table 4. — Experimental variables and test ranges 

Variable Range 

Current density A/dm^. . . 50-150 

Temperature °C . . . 30-50 

Acid concentration pH . . . 4.0-1 .5 

Zn concentration g/L. . . 90-150 

Flow rate m/s . . . 2-5 



The deposits prepared with electrolytes made from the waste 
dusts were evaluated in a similar manner using a factorial design. 
The two parameters, acid and Zn concentration, however, were 
held at 1.5 pH and 90 g/L, respectively. 



STEEL PREPARATION 

The steel sheet was supplied by Inland Steel and was 0.79-mm 
thick AKDQ alloy. As received, it had been sheared to size for 
the flow cells and oiled. The specimens were degreased with sol- 



ELECTROCHEMICAL CORROSION EVALUATION 

Coupons were stamped from the electrogalvanized sheet to fit 
a flat specimen holder and yield a 1-cm^ surface area. The coupons 
were degreased in boiling trichloroethylene, rinsed in ethanol, and 
dried in a filtered airstream. 



60 



The cell used was a commercially available 1-L glass vessel 
with various necks for electrodes, gas inlet and outlets, and ther- 
mometer (7). The counterelectrode was Pt mesh, and the reference 
was a saturated calomel electrode. The medium was analytical rea- 
gent grade (NH4)2S04 of 1 mol/L concentration and pH of 6+0. 1 . 
Prior to sample immersion, the medium was deaerated with oxygen- 
free N, and the gas purge continued throughout the experiments. 
All experiments were carried out at 25° + l ° C. 

Impedance measurements were made at the open circuit poten- 
tial (OCP) using a lock-in amplifier over a frequency range of 10 
to 20,000 Hz having a peak amplitude of 5.15 mV (3.64 mV RMS). 
Measurements from 0.005 to 1 1 Hz were obtained in the time 
domain by an EG&G' fast fourier transform technique. The soft- 
ware performs the experiment and calculates the data points. The 
samples were immersed in the test medium for 1 h before measur- 
ing the electrochemical impedance. 

FORMABILITY TESTS 

Both compression and tension bend tests were used to evalu- 
ate the deposits. The compression samples were bent 180°, with 
the coating on the inside of the bend, and then straightened. Scotch 



brand tape was applied to the distorted area and removed. The rela- 
tive amount of coating powdering was recorded. The tension sam- 
ples were bent 180°, with the coating on the outside of the bend, 
and examined under low magnification to determine if peeling or 
flaking occurred in the deformed area. Then, the specimen was bent 
repeatedly back and forth over a mandrel until the steel fractured. 
The fracture area was examined under low magnification for sepa- 
ration or peeling of the coating. Prying with a sharp knife was used 
to indicate unsatisfactory adhesion by lift off of the coating (8). 
In addition to bend tests, drawing and ball punch deformation 
tests were conducted on coated sheet. Round 92-mm-diam blanks 
were punched from coated sheet and drawn into flanged cups 
approximately 43 mm in diameter and 42 mm deep. The ball punch 
deformation test was conducted in accordance with ASTM E643-78 
(9) procedures. As an added indication of adhesion, a reverse ball 
punch test was also conducted on electrogalvanized sheet. In this 
test, the sheet was placed on the ram of the ductility tester, Zn side 
down, to form a 7.8-mm cone. The specimen was then removed 
and turned over with the top of the cone centered on the ram. The 
cone was then pushed back through the plate for a total of 14 mm. 
The condition of the Zn coating in the deformed area was then exam- 
ined for flaking or peeling. 



RESULTS AND DISCUSSION 



CURRENT EFFICIENCY AND ORIENTATION 

Current efficiency was calculated from the amount of electric 
charge used for a given sample and the weight of the deposit. Three 
specimens were coated for each set of conditions to obtain an aver- 
age value. In the experimental work using the pure zinc sulfate elec- 
trolyte and small flow cell, the current efficiencies ranged from 
95.2 to 99.0 pet. Thirteen out of the sixteen experimental combi- 
nations produced current efficiencies within 1 pet of each other in 
this range. This indicates little difference in the effects of the vari- 
ables over the test ranges in table 4. 

Similar results were obtained with the electrolytes prepared 
from the oxide fume and the EAF dust. The current efficiencies 
were in the same range of upper 90's values indicating no serious 
decrease due to the impurities present, especially the CI. High levels 
of CI did result in increased attack of the Pb-Ag anode producing 
nonconducting surface coatings, which resulted in higher voltages 
to hold the desired current settings. 

The percentage of a particular crystallographic orientation that 
was used in the factorial analysis was arrived at by dividing the 
total count in the X-ray diffraction peak for that crystallographic 
plane by the total count for all planes and multiplying by 100. Table 
5 lists some typical results obtained for selected conditions. For 
the pure Zn electrolyte, the predominant orientations were 002 and 
103. It was noted that current density and Zn concentration played 
only a very minor role in either the orientation or current efficiency 
analysis. 

Deposits produced from oxide fume electrolytes were differ- 
ent than those from the pure Zn electrolyte in that little (002) orien- 
tation was produced under any of the conditions. The majority of 
the deposits were 1 12 and 101 . The same was true for the electro- 
lyte from the EAF dust where there were higher percentages of 
all orientations other than 1 12 and 101 . Examples of typical deposits 
are shown in the SEM photomicrographs in figure 1 . The grain 
size of the deposits produced from waste electrolytes were also 
smaller than those from pure zinc oxide and from the industrial 
sample. 



Preferred orientations of the industrial sample used for com- 
parison were primarily 104, 103, 004, and 002, in order of descend- 
ing percentage. Similar orientations were obtained for deposits 
prepared from pure electrolyte in the large cell. 

ELECTROCHEMICAL CORROSION 

Table 6 contains the ac data (electrochemical impedance spec- 
troscopy) for selected samples from the tests on electrogalvanized 
samples. The values of polarization resistance (Rp) and double- 
layer capacitance (C^l) have been shown in previous research (10) 
to be useful in monitoring the performance of electrogalvanized 



Table 5.— Preferred crystallographic orientation of coated 

deposits prepared from different electrolytes at 500 A/dm^, 

90 g/L Zn, 1.5 pH, 60° C, 2 m/s, percent of total 



Deposit 


















source 


002 


004 


101 


102 


103 


104 


110 


112 


Industrial 


10 


20 








28 


36 








Pure ZnO 


73 


6 





2 


14 


2 








EAF dust 


14 


26 





11 


17 


13 





9 


Oxide fume . . . 








39 


4 


1 





5 


45 



Table 6.— Alternating current electrochemical impedance data 
for electrogalvanized steel in 1M (NH4)2S04 at 25° C after 1 h 



^ Reference to specific products does not imply endorsement by the Bureau of Mines. 







Polarization 


Double-layer 


Deposit 


Plating current 


resistance, 


capacitance, 


source 


density, A/dm^ 


(Rp), ncm2 


(C^l), /iF/cm2 


Industrial 


60 


1,600 


15 


Pure ZnO 


50 


1,625 


17 




150 


800 


28 


EAF dust 


50 


1,200 


13 




150 


600 


20 


Oxide fume .... 


50 


800 


28 




150 


500 


42 



61 





""W^ 



mv^- 




t. 

19 



V.:-^- 



'..^^ ' 






:-.^*> 

..^^'^r^ 



2 

.1 






■is 



^ >" i*» 






Figure 1 .— SEM photomicrographs of electrogalvanized samples. (A) Industrial sample and samples electrogalvanlzed with (B) 
pure ZnO electrolyte, {C) EAF dust electrolyte, and (D) oxide fume electrolyte. 



wire. It demonstrated that the corrosion rate of electrogalvanized 
steel was related to the ac impedance value of the polarization resis- 
tance in deaerated molar (NH4)2S04 under near-neutral conditions. 
The value of the double-layer capacitance was also reported to be 
related to the corrosion rate and to the surface condition of the metal 
coating. The corrosion rate (icorr^) varies inversely as the polari- 
zation resistance. Thus, the larger the value of Rp the lower the 
corrosion rate. Low values of Cjl are usually obtamed when there 
are surface films such as oxides or hydroxides of Zn and when the 
deposit is slowly corroding. Thus, low values indicate low corro- 
sion rates when accompanied by large values of Rp. 



It is seen in examining the electrochemical data in table 6, with 
the preceding in mind, that the corrosion rates are similar for the 
industrial, pure ZnO and EAP dust derived deposits that were plated 
at the lower current density. The deposit from the oxide fiime 
exhibits a higher corrosion rate, which may be the result of the high 
CI level in the electrolyte and the increased grain boundary area 
due to smaller grain size. 

The data for the high-current-density deposits show what would 
amount to an increase in corrosion rate over the low-density plated 
samples. The EAF dust derived and the pure ZnO deposits are simi- 
lar, with a possible advantage to the EAF dust deposit. Again, the 



62 



oxide fume exhibits values indicating higher corrosion rates over 
the pure ZnO and EAF dust deposits. 

Results of salt spray corrosion tests by an independent labora- 
tory on pure ZnO deposits indicated a similar trend. Only current 
density was significant in influencing salt spray results. Deposits 
plated at 50 A/dm^ produced ASTM ratings of 8 to 9.5 (11), while 
those at 150 A/dm^ ranged from 4 to 6. The apparent reason is 
that current density influences nucleation and growth during elec- 
trodeposition and has a pronounced effect on grain size, with more 
active grain boundary material becoming predominant. 

FORMABILITY TESTS 

The results of selected bend tests are shown in table 7. All speci- 
mens bent in compression and straightened exhibited a few small 
cracks. However, only one specimen had some coating material 
removed with the tape. That sample was prepared with pure ZnS04 
at a current density of 150 A/dm^. 

The tension bend tests resulted in small cracks on the indus- 
trial and oxide fume deposits. All others exhibited smooth bends. 
The examination of the coatings, after they were repeatedly bent 
through 180° until failure of the steel, indicated no separation or 
peeling of the coatings. Attempts to pry the coatings loose with a 
knife edge were unsuccessful, indicating good adhesion. 

Microscopic examination of the drawn cup samples showed 
no cracking, flaking, or peeling of the Zn coatings, which indi- 
cates good ductility and adhesion. Examples of the drawn samples 
are shown in figure 2, which illustrates this. 

The ball punch deformation test, used to evaluate the forma- 
bility of sheet materials, was also used on the coated sheet. Sam- 



ples plated from pure electrolyte at 60° C had cup heights and 
maximum loads similar to bare steel and commercial electrogal- 
vanized sheet, while those plated at 35° C, with coarser-grained 
deposits, had lower cup heights and loads. Ball punch tests on sam- 
ples prepared from EAF dust electrolyte indicated good formabil- 
ity as evidenced by cup heights and maximum loads similar to 
industrial samples. No peeling or flaking of the coating occurred 
in the area-of-rupture as shown in figure 3. 

The reverse ball punch specimen shown in figure 4A is from 
a commercially coated sheet. The inner "ring," and just above it, 
is the area of severest deformation. There was no cracking, peel- 
ing, or flaking of the coating in this area. The specimens prepared 
with pure electrolyte were very similar to the commercial speci- 
men. Figure 4fi shows the EAF specimen, which exhibits some 
cracking at the ring. However, it was not possible to peel the coat- 
ing at the cracks using a sharp knife point. This amount of crack- 
ing would not prevent the coating from being acceptable. 



Table 7. — Bend test results on electrogalvanlzed steel sheet 

Deposit Plating current Compression Tsnsion Bend rupture 

source density, A/dm^ Crack Powdering cracks Bends Peeling 

Industrial .... 60 Yes . No Small . 12 No. 

Pure ZnO ... 50 Yes No No ... . 10 No. 

150 Yes. Moderate. No . . . . 9 No. 

EAF dust 50 Yes . No No ... . 9 No. 

150 Yes . No No 10 No. 

Oxide fume.. 50 Yes. No Small. 10 No. 

150 Yes . No No 11 No. 



CONCLUSIONS 



It has been shown that Zn from steel wastes such as EAF dusts 
can be used to successfully electrogalvanize steel sheet. Purifica- 
tion of the leach liquor is necessary, however, to reduce coextracted 
impurities Fe, Cu, Cd, and CI. Electrogalvanizing tests in flow cells 
using pure and waste-derived electrolytes yielded good deposits at 
comparable current efficiencies which were in the 94- to 99-pct 
range. Deposits from waste electrolytes exhibited smaller grain size 
and less basal plane orientation than pure ZnS04 electrolyte. 

Electrochemical corrosion evaluation of coated deposits showed 
that deposits from EAF dust, pure ZnO, and an industrial line have 



similar corrosion rates, while deposits from oxide fume with high 
CI content have higher corrosion rates. Electrochemical data corre- 
late with salt spray corrosion tests in showing that current density 
influences corrosion rate. 

Bend, draw, and ball punch tests on coated steel sheet prepared 
from pure ZnO, EAF dust, and bend tests on oxide fume derived 
deposits indicated that the adhesion and ductility of the coatings 
were as good as those of an industrial sample. 



63 




Figure 2.— Cups drawn from electrogalvanized samples. (A), Industrial sample and cups drawn from samples electrogalvanized 
with (B) pure ZnO electrolyte and (C) EAF dust electrolyte. 



64 



B 



i 





Figure 3.-Specimens from ball punch tests of electrogalvanized samples. (A) Industrial sample and (B) sample from EAF dust 
electrolyte. 



A 



65 



-'i^l 




B 



,^ 



y 



y 



dus^elect^ytr''"'"' ''°'" '''"'' ''" P""^*^ ^^^*^ °" electrcgalvanized samples. W Industrial sample and iB) sampi 



e from EAF 



66 



REFERENCES 



1. Jolly, J. H. Zinc. Sec. in BuMines Mineral Commodity Summaries 
1987, pp. 180-181. 

2. Carrillo, F. V., M. H. Hibpshman, and R. D. Rosenkranz. Recov- 
ery of Secondary Copper and Zinc in the United States. BuMines IC 8622, 
1974, 58 pp. 

3. Krishnan, E. R., and W. F. Kenner. Recovery of Metallic Values 
From Electric Arc Furnace Steelmaking Dusts. Paper in Proceedings of 
Symposium on Iron and Steel Pollution Abatement Technology for 1982. 
EPA Center for Environ. Res. Inf., Research Triangle Park, NC, 1983, 
687 pp. 

4. U.S. Code of Federal Regulations. Title 40— Protection of Environ- 
ment; Chapter I— Environmental Protection Agency; Subchapter I— Solid 
Wastes. Part 261— Identification and Listing of Hazardous Waste; Subpart 
D— List of Hazardous Wastes, July 1, 1983. 

5. Adaviya, T., M. Omura, K. Matsudo, and H. Naemura. Develop- 
ment of Corrosion-Resistant Electrogalvanized Steel. Plat. Surf. Finish., 
v. 68, No. 6, 1979, pp. 96-99. 

6. Dattilo, M., E. R. Cole, Jr., and T. J. O'Keefe. Recycling of Zinc 
Waste for Electrogalvanizing. Conserv. & Recycling, v. 8, No. 3-4, 1985, 
pp. 399-409. 

7. American Society for Testing and Materials. Standard Recommended 
Practice for Standard Reference Method for Making Potentiostatic and Poten- 



tiodynamic Anodic Polarization Measurements. G5-78 in 1982 Annual Book 
of ASTM Standards. Part 10 — Metals — Mechanical, Fracture, and Corro- 
sion Testing; Fatigue; Erosion and Wear; Effect of Temperature. Philadel- 
phia, PA, 1982, pp. 906-916. 

8. . Standard Test Methods for Adhesion of Metallic Coat- 
ings. B571-79 in 1982 Annual Book of ASTM Standards. Part 9, Metallic 
and Inorganic Coatings; Metal Powders, Sintered P/M Structural Parts. 
Philadelphia, PA, 1982, pp. 419-422. 

9. .Standard Methods for Conducting a Ball Punch Defor- 
mation Test for Metallic Sheet Material. E643-78 in 1982 Annual Book 
of ASTM Standards. Part 10, Metals — Mechanical, Fracture, and Corro- 
sion Testing; Fatigue, Erosion and Wear; Effect of Temperature. Philadel- 
phia, PA, 1982, pp. 758-761. 

10. Dattilo, M. The Use of AC Impedance to Determine the Corrosion 
Rate of Electrogalvanized Steel. Mater. Perf., v. 35, No. 11, Nov. 1986, 
pp. 18-22. 

1 1 . American Society for Testing and Materials. Standard Recommended 
Practice for Rating of Electroplated Panels Subjected to Atmospheric 
Exposure. B537-70 (Reapproved 1981) in 1982 Annual Book of ASTM 
Standards. Part 9, Metallic and Inorganic Coatings; Metal Powders, Sin- 
tered P/M Structural Parts. Philadelphia, PA, 1982, pp. 364-374. 



67 



ECONOMIC EVALUATION OF A TECHNIQUE TO PELLETIZE FLUE DUST 
AND OTHER WASTE FROM THE MANUFACTURE OF STAINLESS STEEL 



By Joan H. SchwieM 



ABSTRACT 

Contained in this paper is an economic evaluation of a method to pelletize flue dust and other 
wastes from the manufacture of stainless steel. Bureau of Mines personnel have demonstrated that 
the alloying elements contained in this waste can be recovered if these pellets are used to replace 
a portion of the scrap fed to an electric arc furnace producing stainless steel. This evaluation con- 
siders two pellet drying options— heat dried and air dried. 

The fixed capital cost for a plant addition required to produce 15 st of pellets per day is esti- 
mated to be about $974,000 for the heat-dried option and $560,000 for the air-dried option, based 
on second quarter 1987 equipment costs. The estimated annual operating cost per short ton of pellets 
to pelletize these wastes is approximately $117/st with the heat-dried option and $40/ st with the 
air-dried option. The value of the pellets, based on the value of the contained alloying elements 
as ferroalloys, is estimated to be about $313/st. This indicates that the proposed pelletizing tech- 
nique has economic potential. 



INTRODUCTION 



The manufacture of stainless steel results in the production of 
several wastes such as grinding swarf, mill scale, and flue dust from 
electric arc and argon-oxygen decarbonization furnaces. In an effort 
to recycle these wastes, the Bureau of Mines has investigated a tech- 
nique to pelletize the wastes recovered from an electric-arc fur- 
nace. The original research^ centered on smelting these pellets 
separately then recycling the recovered metal alloy as a replace- 
ment for the scrap portion of a stainless steel production heat. 

Additional research has shown that it is also feasible to replace 
a portion of the scrap charge for stainless steel production heats 



with the pelletized waste. The pellets can compose from 10 to 20 
pet of the total charge to the furnace. This enables recycling of the 
waste without the intermediate smelting stage. In this manner, it 
is expected that the alloying elements present in the waste can be 
recycled at minimum cost to the manufacturer. 

This evaluation is the latest in a series of evaluations that has 
been used to help guide the research. The research work is described 
in another paper in this Information Circular, titled "Recycling of 
Stainless Steelmaking Dusts and Other Wastes," by L. A. Neu- 
meier and M. J. Adam. 



PROCESS DESCRIPTION 



A plant addition has been designed to produce 15 st/d of pellets 
from stainless steel wastes, operating one shift per day, 5 days per 
week. 

Wastes that are assumed to be available for the production of 
the pellets are grinding swarf, mill scale, and flue dusts from elec- 
tric arc and argon-oxygen decarbonization furnaces. The flue dusts, 
mill scale, and grinding swarf are stored in open piles and moved 



' Cost evaluation program assistant. Process Evaluation, Bureau of Mines, 
Washington, DC. 

2 Powell, H. E., W. M. Dressel, and R. L. Crosby. Converting Stainless Steel Fur- 
nace Flue Dust and Wastes to a Recyclable Alloy. BuMines RI 8039, 1975, 24 pp. 



daily to storage bins with a front-end loader. Mill scale is screened 
to separate it into three fractions — plus % in, minus % in plus 35 
mesh, and minus 35 mesh. The plus %-in fraction is large enough 
to be recycled directly and is not included in the pellets. It is con- 
veyed to a storage bin untU needed in the furnace. Minus %-in plus 
35-mesh mill scale is conveyed to a ball mill and crushed. The ball 
mill product is returned to the screen. Minus 35-mesh mill scale 
from the screen is conveyed to a storage bin. 

In a zig-zag mixer, the mill scale is combined with grinding 
swarf and flue dusts from electric arc and argon-oxygen decarboni- 
zation furnaces. Coke breeze and portland cement are also added 
to the mixer. The coke breeze is required to reduce the oxidized 



68 



portion of the waste when the pellets are smelted. Portland cement 
is added as a binding agent. 

The combined wastes are then conveyed to a balling drum for 
pelletization. The resulting green pellets, containing about 12 pet 
moisture, are dried by one of two methods, the heat-dried option 
using truck dryers or the air-dried option using the same type dry- 
ing trucks, but without a dryer. In the heat-dried option, the pellets 
are air dried for 24 h before being completely dried in a heated 
dryer. Using the air-dried option, the pellets are dried in the open 
for several days. Either drying method produces pellets suitable 
for feeding to an electric arc furnace. The only difference in the 
plant designs for the two options is the addition of two truck dryers 
for the heat-dried option. 

The evaluation is based on pellets produced from the materials 
presented in table 1 . The composition of the dry pellets is shown 
in table 2. There is no reason to believe that the raw material ratios 
used in this study could not be changed to permit the use of vary- 
ing quantities of each waste. 



Table 1 .—Materials used to make up the pellets, percent 

Argon-oxygen decarbonization furnace flue dust 13.0 

Electric arc furnace flue dust 17.4 

Grinding swarf 39.1 

Mill scale 17.4 

Portland cement 4.4 

Coke breeze 8.7 

Total 100.0 

Table 2.— Pellet composition (dry basis), percent 

Cfiromium 9.5 

Nickel 3.96 

Molybdenum .84 

Manganese 2.0 

Iron 41 .8 

Carbon 11.8 

Silicon 3.5 

Other 26.6 

Total 100.00 



ECONOMICS 



The intent of an economic evaluation is to present capital and 
operating cost estimates of a commercial-size plant. In the prepara- 
tion of any economic evaluation, it is necessary to make many 
assumptions. In general, the assumptions that are made are expected 
either to apply to the majority of the potential plants or to have 
only a small effect on the process capital and operating costs. An 
example of such an assumption is that the plant operates one shift 
per day, 5 days per week. 

If an assumption would be necessary that may not apply to a 
majority of plants or may have a major effect on capital or operat- 
ing costs, then it is generally not included in the evaluation. An 
example of such an exclusion is that land cost and pond construc- 
tion costs have not been included in the capital or operating cost 
estimates. When an assumption has been made or deliberately 
excluded, this fact is documented. 

A detailed description of the estimating techniques used in this 
evaluation has been published.^ 

CAPITAL COSTS 

The capital cost estimate is of the general type called a study 
estimate by Weaver and Bauman." This type of estimate, prepared 
from a flowsheet and a minimum of equipment data, can be expected 
to be within 30 pet of the actual cost. Equipment costs are from 
informal cost quotations from equipment manufacturers and from 
capacity-cost data. The costs of the major items of equipment and 
their accessories are tabulated in the appendix to this paper. 

The estimated fixed capital costs for plant additions capable 
of producing 15 st/d of pellets, on a second quarter 1987 basis (Mar- 
shall and Swift (M and S) index of 808.0), are approximately 
$974,000 for the heat-dried option and $560,000 for the air-dried 
option, as shown in table 3. Because this is a plant addition, the 
cost to hook up to existing plant facilities and utilities is estimated 
as 2 pet of the total section costs. 

Factors for piping, etc., except for the electrical factor, are 
assigned to each section, using as a basis the effect fluids, solids, 
or a combination of fluids and solids may have on the process equip- 
ment. The electrical factor is based on the motor horsepower 



' Peters, F. A. Economic Evaluation Methodology. BuMines IC 9147, 1987, 21 pp. 

' Weaver, J. B., and H. C. Bauman. Cost anci Profitability Estimation. Sec. 25 
in Perry's Chemical Engineers's Handbook, ed. by R. H. Perry and C. H. Chilton. 
McGraw-Hill, 5th ed., 1973, p. 47. 



requirements for each section. A factor of 10 pet, referred to as 
miscellaneous, is added to each section to cover minor equipment 
and construction costs that are not shown with the equipment listed. 

For each section, the field indirect cost, which covers field 
supervision, inspection, temporary construction, equipment rental, 
and payroll overhead, is estimated at 10 pet of the direct cost. 
Engineering cost is estimated at 10 pet, and administration and over- 
head cost is estimated at 5 pet of the construction cost. A contin- 
gency allowance of 10 pet and a contractor's fee of 5 pet are included 
in the section costs. 

The costs of plant facilities and plant utilities are estimated as 
2 pet each of the total process section costs and include the same 
field indirect costs, engineering, administration and overhead, con- 
tingency allowance, and contractor's fee as are included in the sec- 
tion costs. Included under plant facilities are the cost of material 
and labor for auxiliary buildings such as offices, shops, laborato- 
ries, and cafeterias, and the cost of nonprocess equipment such as 
office furniture, and safety, shop, and laboratory equipment. Also 

Table 3.— Estimated capital cost,^ heat-dried and air-dried options 



Heat dried 



Air dried 



$142,000 
357,600 



Fixed capital: 

Mill-scale preparation section $142,000 

Mixing and pelletization section 726,900 

Subtotal 

Plant facilities, 2 pet of above subtotal. 
Plant utilities, 2 pet of above subtotal . . 

Total plant cost 

Land cost 

Subtotal 

Interest during construction period 

Fixed capital cost 

Working capital: 
Raw material and supplies inventory. . . 

Product and in-process inventory 

Accounts receivable 

Available casfi 

Working capital cost 

Capitalized startup cost 

Subtotal 

Total capital cost 1 ,096,300 



868,900 
17,400 
17,400 


499,600 
10,000 
10,000 


903,700 



519.600 



903,700 
70,700 


519,600 
40,400 


974,400 


560,000 


2,400 
39,600 
39,600 
30,600 


2,200 
14,800 
14,800 

9,500 


112,200 
9,700 


41,300 
5,600 


121,900 


46,900 



606,900 



^ Basis: M and S equipment cost index of 808.0. 



69 



included are labor and material costs for site preparation such as 
clearing, grading, drainage, roads, and fences. The costs of water, 
power, and steam distribution systems are included under plant 
utilities. 

Working capital is defined as the funds in addition to fixed cap- 
ital, land investment, and startup costs that must be provided to 
operate the plant. Working capital, also shown in table 3, is esti- 
mated from the following items: (1) Raw material and supplies 
inventory (cost of raw material and operating supplies for 30 days), 
(2) product and in-process inventory (total operating cost for 30 
days), (3) accounts receivable (total operating cost for 30 days), 
and (4) available cash (direct expenses for 30 days). 

Capitalized startup costs, are estimated as 1 pet of the fixed cap- 
ital costs, and are shown in table 3. The cost of land is not included 
in this estimate. 



OPERATING COSTS 

The estimated annual operating cost is based on the average 
of 260 days of operation per year, one shift per day, over the life 
of the plant. The operating costs are divided into direct, indirect, 
and fixed costs. 



Direct costs include raw materials, utilities, direct labor, plant 
maintenance, payroll overhead, and operating supplies. Raw 
materials and utility requirements per short ton of pellets are shown 
in the appendix. The shipping charge must be added to the cost 
of the raw material because the plant location has not been selected. 
Payroll overhead, estimated as 35 pet of direct labor and main- 
tenance labor, includes vacation, sick leave, social security, and 
fringe benefits. 

Plant maintenance is separately estimated for each piece of 
equipment and for the buildings, electrical system, piping, plant 
utility distribution systems, and plant facilities. 

The indirect costs include the expenses of control laboratories, 
accounting, plant protection and safety, plant administration, mar- 
keting, and company overhead. These costs are estimated as 40 pet 
of the direct labor and maintenance costs. Research and overall com- 
pany administrative costs outside the plant are not included. 

Fixed costs include the cost of taxes (excluding income taxes), 
insurance, and depreciation. Depreciation is based on a straight- 
line, 20-yr period. 

The net operating cost per short ton of pellets is $1 17 for the 
heat-dried option; and $40 for the air-dried option. These costs are 
presented in table 4. Included in this cost is a credit of 1 cent per 
pound ($20/st) for the reduced landfill requirements. If the dusts 



Table 4.— Estimated annual operating cost, heat-dried and air-dried options 



Heat-dried costs 



Annual 



Per St 
pellets 



Air-dried costs 



Annual 



Per St 
pellets 



Direct cost: 
Raw materials: 

Portland cement at $60/st 

Coke breeze at $32/st 

Argon-oxygen decarbonization dust at $0.00/st . 

Electric arc furnace flue dust at $0.00/st 

Grinding swarf at $0.00/st 

Mill scale at $0.00/st 

Replacement balls for grinding at $0.27/lb .... 
Total 

Utilities: 

Electric power at $0.05/kW h 

Process water at $0.25/Mgal 

Natural gas at $5.25/MMBtu 

Total 

Direct labor: 

Labor at $10.50/h 

Supervision, 20 pet of labor 

Total 

Plant maintenance: 

Labor 

Supervision, 20 pet of maintenance labor 

Materials 

Total 

Payroll overhead, 35 pet of above payroll 

Operating supplies, 20 pet of plant maintenaee . . 
Total direct cost 

Indirect cost, 40 pet of direct labor and maintenaee 
Fixed cost: 

Taxes, 1 pet of total plant cost 

Insurance, 1 pet of total plant cost 

Depreciation, 20-yr life 

Total operating cost 

Credit: 

Reduced landfill requirements at $0.01 /lb 

Net operating cost 



$10,300 
13,400 




100 



23,800 



2,600 

100 

202,000 



204,700 



65,500 
13,100 



78,600 



12,400 

2.500 

12,400 



27,300 



32,700 
5,500 



372,600 



42,400 

9,000 

9,000 

48,700 



481,700 



23,700 



458,000 



$2.64 
3.44 
.00 
.00 
.00 
.00 
.03 



$10,300 
13,400 




100 



6.11 



23,800 



,67 

.03 

51.79 



900 
100 
NAP 



52.49 



1,000 



16.79 
3.36 



43,700 
8,700 



20.15 



52,400 



3.18 

.64 

3.18 



6,400 
1,300 
6,400 



7.00 



14,100 



8.38 
1.41 



21 ,000 
2,800 



95.54 



115,100 



10.87 

2.31 

2.31 

12.49 



123.52 



6.08 



117.44 



26,600 

5,200 

5,200 

28,000 



180,100 



23,700 



156,400 



$2.64 
3.44 
.00 
.00 
.00 
.00 
.03 



6.11 



.23 

.03 

NAP 



.26 



11.21 
2.23 



13.44 



1.64 

.33 

1.64 



3.61 



5.38 
.72 



29.52 



6.82 

1.33 
1.33 
7.18 



46.18 



6.08 



40.10 



NAP Not applicable. 



70 



are declared as hazardous waste, the disposal cost would be at least 
$2(X)/st. The grinding swarf and mill scale may have a market value; 
but a cost for them is not included in the operating cost. 

PRODUCT VALUE 

To estimate the value of the pellets, it is assumed that the con- 
tained alloying elements will have a value equal to their price as 
a ferroalloy or metal. These values, per pound are, nickel, $2.10; 
chromium as ferrochrome, $0.42; molybdenum, $3.20; and man- 
ganese as ferromanganese, $0.33. Based on the composition listed 



in table 2, the value of the pellets is about $313/st. Based on the 
chromium and nickel values alone, the pellet value is about $246/st. 
In either case the value of the metal contained in the pelletized waste 
is much higher than the cost to pelletize it. 

It should be noted that the pellet value listed in the preceding 
paragraph is only an estimate and at best will only be representa- 
tive of pellets with the same composition. For a particular loca- 
tion, with its own unique wastes, the value of pellets will vary 
significantly from the value presented. It is expected, however, that 
the values given are sufficiently representative to allow anyone 
interested in the Bureau's recycling technique to make a decision 
as to whether additional consideration is warranted. 



TECHNICAL EVALUATION 



The technique to pelletize the stainless steel waste, as presented 
in this paper, utilizes standard agglomeration techniques and should 
present no problems in scale-up to a commercial size. Pellets 
produced at the Bureau's Rolla (MO) Research Center were used 
to replace part of the scrap charge to an electric arc furnace at a 
commercial stainless steel manufacturer and were successfully 
smelted. The results of this testing can be obtained from the research 
personnel at the Rolla Research Center. It appears, therefore, that 
the proposed recycling technique has been sufficiently developed 
to allow serious consideration of it for adaptation on a commercial 
scale. 

The use of the proposed process has two potential advantages. 
First, the alloying elements previously lost in the wastes will be 



recycled, which will lower the overall operating cost for the stain- 
less steel manufacturer. Also, because chromium and nickel are 
almost totally imported, their recycle will reduce the U.S. depen- 
dence on these imports. The second advantage will be a reduction 
in the landfill requirements of the stainless steel manufacturer. 
Processing the wastes can only increase the environmental accept- 
ability of a plant. 

Wastes produced by a stainless steel plant will vary from plant 
to plant as well as from day to day. This is due to the variety of 
alloys produced and to variations in equipment and procedures. The 
Bureau's recycling technique, however, should be applicable to any 
fine stainless steel manufacturing waste. 



APPENDIX.— HEAT-DRIED AND AIR-DRIED OPTIONS 



Table A-1 . — Raw material and utility requirements per short ton of pellets 



71 



Heat dried 



Air dried 



Raw materials: 

Portland cement st . , 

Coke breeze st . , 

Argon-oxygen decarbonization dust st . 

Electric arc furnace flue dust st . 

Grinding swarf st . 

Mill scale st . 

Replacement balls for grinding lb . 

Utilities: 

Electric power kWh . 

Process water Mgal . 

Natural gas MMBtu . 



0.044 


0.044 


,107 


.107 


.130 


.130 


.174 


.174 


.391 


.391 


.174 


.174 


.067 


.067 


3.467 


4.533 


.133 


.133 


9.867 


Nap 



Nap Not applicable. 



Table A-2.— Equipment cost summary, mill scale preparation section, heat-dried and air-dried options 

Item Equipment Labor Total 

Mill-scale hopper $800 $300 $1 ,100 

Belt conveyor 7,700 1 ,300 9,000 

Vibrating screen 10,700 1,400 12,100 

Bucket conveyor 1 ,400 400 1 ,800 

Storage bin 1 ,800 700 2,500 

Bucket elevator 3,000 900 3,900 

Belt conveyor 7,000 1 ,200 8,200 

Ball mill 3,500 200 3,700 

Mill-scale storage bin 300 100 400 

Total 36,200 6,500 42,700 

Total equipment cost x factor indicated: 

Foundations, x 0.695 25,200 

Structures, x 0.080 2,900 

Instrumentation, x 0.050 1,800 

Electrical, x 0,362 13,100 

Piping, X 0.200 7,200 

Painting, x 0.020 700 

Miscellaneous, x 0,100 3,600 

Total 54,500 

Total direct cost 97,200 

Field indirect, 10 pet of total direct cost 9,700 

Total construction cost 1 06,900 

Engineering, 10 pet of total construction cost 10,700 

Administration and overhead, 5 pet of total construction cost 5,300 

Subtotal 122,900 

Contingency, 10 pet of above subtotal 12,300 

Subtotal 135,200 

Contractor's fee, 5 pet of above subtotal 6,800 

Section cost 1 42,000 



^ Basis: M and S equipment cost index of 808,0. 



72 



Table A-3.— Equipment cost summary, mixing and pelletization section, heat-dried option 

Item EquipmenV Labor Total 

Argon-oxygen decarbonization dust storage bin $2,000 $700 $2,700 

Argon-oxygen decarbonization dust feeder 1,000 100 1,100 

Electric arc furnace flue dust storage bin 2,000 700 2,700 

Electric arc furnace flue dust feeder 1,000 100 1,100 

Grinding swarf storage bin 2,000 700 2,700 

Grinding swarf feeder 1,000 100 1,100 

Portland cement storage bin 4,200 1 ,200 5,400 

Portland cement feeder 1 ,000 200 1 ,200 

Coke breeze storage bin 3,000 1 ,200 4,200 

Coke breeze feeder 1,000 100 1,100 

Belt conveyor from storage 10,000 2,100 12,100 

Mixer (zig-zag) 1,300 100 1,400 

Belt conveyor 7,700 1,400 9,100 

Pelletizer 48,000 600 48,600 

Belt conveyor 5,800 900 6,700 

6-truck dryer2 68,500 2,400 70,900 

9-truck dryer2 97,000 2,600 99,600 

Pellet storage bin 2,600 1 ,000 3,600 

Bucket conveyor 6,200 1,900 8,100 

Total 265,300 18,100 283,400 

Total equipment cost x factor indicated: 

Foundations, x 0.228 60,400 

Structures, x 0.080 21 ,200 

Insulation, x 0.020 5,300 

Instrumentation, x 0.050 13,300 

Electrical, x 0.109 29,000 

Piping, X 0.200 53,100 

Painting, x 0.020 5,300 

Miscellaneous, x 0.100 26,500 

Total 214,100 

Total direct cost 497,500 

Field indirect, 10 pet of total direct cost 49,800 

Total construction cost 547,300 

Engineering, 10 pet of total construction cost 54,700 

Administration and overhead, 5 pet of total construction cost 27,400 

Subtotal 629,400 

Contingency, 10 pet of above subtotal 62,900 

Subtotal 692,300 

Contractor's fee, 5 pet of above subtotal 34,600 

Section cost 726,900 



1 Basis: M and S equipment cost index of 808.0. 

2 Includes cost of truck. 



73 



Table A-4.— Equipment cost summary, mixing and pelletization section, air-dried option 

Item EquipmenV Labor Total 

Argon-oxygen decarbonization dust storage bin $2,000 $700 $2,700 

Argon-oxygen decarbonization dust feeder 1,000 100 1,100 

Electric arc furnace flue dust storage bin 2,000 700 2,700 

Electric arc furnace flue dust feeder 1,000 100 1,100 

Grinding swarf storage bin 2,000 700 2,700 

Grinding swarf feeder 1,000 100 1,100 

Portland cement storage bin 4,200 1 ,200 5,400 

Portland cement feeder 1 ,000 200 1 ,200 

Coke breeze storage bin 3,000 1 ,200 4,200 

Coke breeze feeder 1 ,000 100 1 ,100 

Belt conveyor from storage 10,000 2,100 12,100 

Mixer (zig-zag) 1 ,300 100 1 ,400 

Belt conveyor 7,700 1,400 9,100 

Pelletizer 48,000 600 48,600 

Belt conveyor 5,800 900 6,700 

Pellet storage bin 2,600 1 ,000 3,600 

Bucket conveyor 6,200 1,900 8,100 

Total 99,800 13,100 112,000 

Drying trucks 16,500 

Total equipment cost x factor indicated: 

Foundations, x 0.463 46,200 

Structures, x 0.080 8,000 

Instrumentation, x 0.050 5,000 

Electrical, x 0.241 24,100 

Piping, x 0.200 20,000 

Painting, x 0.020 2,000 

Miscellaneous, x 0.100 10,000 

Total 1 15,300 



Basis: M and S equipment cost index of 808.0. 



Total direct cost 244,700 

Field indirect, 1 pet of total direct cost 24,500 

Total construction cost 269,200 

Engineering, 10 pet of total construction cost 26,900 

Administration and overfiead, 5 pet of total construction cost 13,500 

Subtotal 309,600 

Contingency, 1 pet of above subtotal 31 ,000 

Subtotal 340,600 

Contractor's fee, 5 pet of above subtotal 17,000 

Section cost 357,600 



U.S. GOVERNMENT PRINTING OFFICE: 1988 - 547000/80,059 INT.-BU.OF MINES,PGH.,PA. 28759 



380 



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U.S. Dapartmant of th« Intarior 
BwMM of MifiM-Prod. and Dtotr. 
Cochrans Mill Road 
P.O. Box 18070 
Pittsburgh, Pa. 15236 



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